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LABORATORY 


technical library 
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armed forces 

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SUMMARY TECHNICAL REPORT 
OF THE 

NATIONAL DEFENSE RESEARCH COMMITTEE 


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Manuscript and illustrations for this volume were prepared for 
publication by the Summary Reports Group of the Columbia 
University Division of War Research under contract OEMsr-1131 
with the Office of Scientific Research and Development. This vol¬ 
ume was printed and bound by the Columbia University Press. 

Distribution of the Summary Technical Report of NDRC has been 
made by the War and Navy Departments. Inquiries concerning 
the availability and distribution of the Summary Technical Report 
volumes and microfilmed and other reference material should be 
addressed to the War Department Library, Room 1A-522, The 
Pentagon, Washington 25, D. C., or to the Office of Naval Re¬ 
search, Navy Department, Attention: Reports and Documents 
Section, Washington 25, D. C. 

Copy No. 

157 


This volume, like the seventy others of the Summary Technical 
Report of NDRC, has been written, edited, and printed under 
great pressure. Inevitably there are errors which have slipped past 
Division readers and proofreaders. There may be errors of fact not 
known at time of printing. The author has not been able to follow 
through his writing to the final page proof. 

Please report errors to: 

JOINT RESEARCH AND DEVELOPMENT BOARD 
PROGRAMS DIVISION (STR ERRATA) 

WASHINGTON 25, D.C. 

A master errata sheet will be compiled from these reports and sent 

to recipients of the volume. Yfeur help will make this book more 

*•. 

useful to other readers and will be of great value in preparing any 
revisions. 



UNCLASSIFIED 

SUMMARY TECHNICAL REPORT OF DIVISION 1, NDRC 

| 

VOLUME 1 


HYPERVELOCITY GUNS 
AND THE CONTROL 
OF GUN EROSION 


OFFICE OF SCIENTIFIC RESEARCH AND DEVELOPMENT 
VANNEVAR BUSH, DIRECTOR 


NATIONAL DEFENSE RESEARCH COMMITTEE 


B. 

CONANT, CHAIRMAN 

r r 



\ I 
A v lV 

D 

IVISION 1 

A A 

H 

. ADAMS, CHIEF 

• <> 



U L t L 


LL 






^33 


WASHINGTON, D.C., 1946 




NATIONAL DEFENSE RESEARCH COMMITTEE 


James B. Conant, Chairman 
Richard C. Tolman, Vice Chairman 
Roger Adams Army Representative 1 

Frank B. Jewett Navy Representative 2 

Karl T. Compton Commissioner of Patents 3 

Irvin Stewart, Executive Secretary 



1 Army Representatives in order of service: 


Maj. Gen. G. V. Strong 
Maj. Gen. R. C. Moore 
Maj. Gen. C. C. Williams 
Brig. Gen. W. A. Wood, Jr. 


Col. L. A. Denson 
Col. P. R. Faymonville 
Brig. Gen. E. A. Regnier 
Col. M. M. Irvine 


Col. E. A. Routheau 


2 Navy representatives in order of service: 

Rear Adm. H. G. Bowen Rear A dm. J. A. Furer 

Capt. Lybrand P. Smith Rear Adm. A. H. Van Keuren 

Commodore H. A. Schade 
3 Commissioners of Patents in order of service: 

Conway P. Coe Casper W. Ooms 


NOTES ON THE ORGANIZATION OF NDRC 


The duties of the National Defense Research Committee were 
(1) to recommend to the Director of OSRD suitable projects 
and research programs on the instrumentalities of w r arfare, 
together with contract facilities for carrying out these projects 
and programs, and (2) to administer the technical and scienti¬ 
fic work of the contracts. More specifically, NDRC functioned 
by initiating research projects on requests from the Army or 
the Navy, or on requests from an allied government trans¬ 
mitted through the Liaison Office of OSRD, or on its own con¬ 
sidered initiative as a result of the experience of its members. 
Proposals prepared by the Division, Panel, or Committee for 
research contracts for performance of the work involved in 
such projects were first reviewed by NDRC, and if approved, 
recommended to the Director of OSRD. Upon approval of a 
proposal by the Director, a contract permitting maximum flex¬ 
ibility of scientific effort was arranged. The business aspects of 
the contract, ^infcluding such matters as materials, clearances, 
vouchers, patents, priorities, legal matters, and administration 
of patent matters were handled by the Executive Secretary of 
OSRD. 

Originally NDRC administered its work through five divi¬ 
sions, each headed by one of the NDRC members. These were: 

Division A—Armor and Ordnance 

Division B—Bombs, Fuels, Gases, & Chemical Problems 
Division C—Communication and Transportation 
Division D—Detection, Controls, and Instruments 
Division E—Patents and Inventions 


In a reorganization in the fall of 1942, twenty-three adminis¬ 
trative divisions, panels, or committees were created, each with 
a chief selected on the basis of his outstanding work in the par¬ 
ticular field. The NDRC members then became* & reviewing 
and advisory group to the Director of OSRD. The'final organ¬ 
ization was as follows: 

Division 1—Ballistic Research 

Division 2—Effects of Impact and Explosion 

Division 3—Rocket Ordnance 

Division 4—Ordnance Accessories 

Division 5—New Missiles 

Division 6—Sub-Surface Warfare 

Division 7—Fire Control 

Division 8—Explosives 

Division 9—Chemistry 

Division 10—Absorbents and Aerosols 

Division 11—Chemical Engineering 

Division 12—Transportation 

Division 13—Electrical Communication 

Division 14—Radar 

Division 15—Radio Coordination 

Division 16—Optics and Camouflage 

Division 17—Physics 

Division 18—War Metallurgy 

Division 19—Miscellaneous 

Applied Mathematics Panel 

Applied Psychology Panel 

Committee on Propagation 

Tropical Deterioration Administrative Committee 


IV 


CONFIDENTIAL 



NDRC FOREWORD 


As events of the years preceding 1940 revealed 
more and more clearly the seriousness of the 
world situation, many scientists in this country came 
to realize the need of organizing scientific research for 
service in a national emergency. Recommendations 
which they made to the White House were given care¬ 
ful and sympathetic attention, and as a result the Na¬ 
tional Defense Research Committee (NDRC) was 
formed by Executive Order of the President in the 
summer of 1940. The members of NDRC, appointed 
by the President, were instructed to supplement the 
work of the Army and the Navy in the development 
of the instrumentalities of war. A year later, upon the 
establishment of the Office of Scient ific Research and 
Development (OSRD), NDRC became one of its 
units. 

The Summary Technical Report of NDRC is a 
conscientious effort on the part of NDRC to summar¬ 
ize and evaluate its work and to present it in a useful 
and permanent form. It comprises some seventy vol¬ 
umes broken into groups corresponding to the NDRC 
Divisions, Panels, and Committees. 

The Summary Technical Report of each Division, 
Panel, or Committee is an integral survey of the work 
of that group. The first volume of each group’s report 
contains a summary of the report, stating the prob¬ 
lems presented and the philosophy of attacking them, 
and summarizing the results of the research, develop¬ 
ment, and training activities undertaken. Some vol¬ 
umes may be “state of the art” treatises covering 
subjects to which various research groups have con¬ 
tributed information. Others may contain descrip¬ 
tions of devices developed in the laboratories. A mas¬ 
ter index of all these divisional, panel, and committee 
reports, which together constitute the Summary 
Technical Report of NDRC, is contained in a sep¬ 
arate volume, that also includes the index of a micro¬ 
film record of pertinent technical laboratory reports 
and reference material. 

Some of the NDRC-sponsored researches which 
had been declassified by the end of 1945 were of 
sufficient popular interest that it was found desirable 
to report them in the form of monographs, such as 
the series on radar by Division 14 and the monograph 
on sampling inspection by the Applied Mathematics 
Panel. Since the material treated in them is not du¬ 
plicated in the Summary Technical Report of NDRC, 
the monographs are an important part of the story 
of these aspects of NDRC research. 


In contrast to the information on radar, which is 
of wide-spread interest and much of which is released 
to the public, the research on subsurface warfare is 
largely classified and is of general interest to a more 
restricted group. As a consequence, the report of 
Division 6 is found almost entirely in its Summary 
Technical Report, which runs to over 20 volumes. 
The extent of the work of a division cannot therefore 
be judged solely by the number of volumes devoted 
to it in the Summary Technical Report of NDRC: 
account must be taken of the monographs and avail¬ 
able reports published elsewhere. 

Division 1, under the leadership of its Chief, L. H. 
Adams, conducted a program of research in the field 
of hypervelocity guns. Its main responsibility was to 
study the effects of gun erosion and the erosive qual¬ 
ities of the common propellants, in order that means 
could be found for prolonging the life of guns firing 
at muzzle velocities higher than the pre-war maxi¬ 
mum of 3,000 feet per second. 

The contractor of the Division developed several 
highly erosion-resistant materials, and means were 
found for the preparation of these alloys in forms 
suitable for use as bore-surface material. Stellite, one 
of these alloys, when applied to caliber .50 machine- 
gun barrels as a short breech liner, increased their 
life several fold under severe conditions, especially 
when the bore surface ahead of the liner was chrom¬ 
ium plated. As the war came to an end, an experimen¬ 
tal 90-mm gun (to fire at a muzzle velocity of at 
least 4,000 fps) was being developed, and further 
work on it was turned over to the Army, while the 
Navy was able to take over the plans for the fabrica¬ 
tion of a molybdenum liner for its new 3-inch/70- 
caliber gun. 

The Summary Technical Report of Division 1, 
prepared under the direction of the Division Chief, 
and authorized by him for publication, is a record of 
this work, and a comprehensive review of the state 
of the art. It also stands as a testimonial to the energy 
and resourcefulness of the personnel of the Division, 
whose contributions to the war effort are worthy of 
grateful appreciation. 

Vannevar Bush, Director 
Office of Scientific Research and Development 

J. B. Conant, Chairman 
National Defense Research Committee 


CONFIDENTIAL 


v 






FOREWORD 


A four-year program of research and develop¬ 
ment in which experienced investigators from 
leading universities, from other academic institu¬ 
tions, and from commercial organizations with com¬ 
petent research and engineering departments work 
together on a series of closely integrated problems is 
almost certain to uncover a variety of interesting 
facts and generalizations. In the case of a broad as¬ 
signment, like that of Division 1, NDRC, involving 
a number of fields in chemistry, physics, metallurgy, 
mathematics, and engineering, it is especially desira¬ 
ble for the new knowledge to be consolidated in the 
form of a Summary Technical Report, as specified by 
OSRD, for the benefit of those who in the future 
may wish to pursue further the problems that were 
attacked. Although the findings of Division 1 are 
covered in great detail by a series of 144 formal re¬ 
ports on various subjects, in addition to numerous 
interim reports submitted by the 29 different con¬ 
tractors, it was decided by Division 1 that its Sum¬ 
mary Technical Report should be more than merely 
a collection or a condensation of previous reports. 
Instead, as is brought out in the Editor’s preface, it 
deals systematically, although briefly, with the sub¬ 
ject of hypervelocity and the control of gun erosion 
in relation to work of the Division. It is intended 
that by the selection of results and conclusions, as 
well as by the sequence of topics, the Report should 
tell a story of an interesting adventure in scientific 
collaboration. 

In formulating the program and in supervising the 
work of the various groups, it was deemed essential 
that a well-rounded series of investigations, includ¬ 
ing research as well as development and testing, be 
set up. It was felt that even in wartime, fundamen¬ 
tal research is essential for real progress. No small 
part of whatever success Division 1 may have at¬ 
tained in a field requiring fresh ideas and novel pro¬ 
cedures was largely a result of the attitude of NDRC 
toward research. In conformity with the fixed policy 
of that agency of the Government with reference to 
the handling of contracts, particular care was exer¬ 
cised in Division 1 not to have Government officials 
substitute their own ideas for those of the skilled 
scientists in the associated research groups. Rather, 
the investigators, engaged as they were in exploring 
the unknown, were given the utmost freedom of ac¬ 


tion consistent with focusing their efforts on the 
central problem. 

The results of those efforts are outlined in this vol¬ 
ume. It may be seen that the accomplishments are 
mainly of two kinds: First, the obtaining of a better 
understanding of gun erosion and the means of com¬ 
batting it; and second, the development and testing 
of specific methods for the solution of the hypervelo¬ 
city problem. As a by-product of the main investiga¬ 
tion, there was obtained a notable improvement in 
machine gun barrels—the need for which was not 
visualized when the original program was formu¬ 
lated. 

To all those who collaborated or otherwise aided 
in this undertaking, the Division Chief is deeply in¬ 
debted: the contractors’ employees who worked to¬ 
gether energetically and harmoniously in the com¬ 
mon cause; the Division staff members who per¬ 
formed patiently and effectively a variety of difficult 
tasks; and the technical representatives of our Bri¬ 
tish and Canadian allies who gave encouragement 
and valuable advice. It is a pleasant duty to record 
also the debt of gratitude that Division 1 owes to 
those individuals in the Navy Department and the 
War Department, acting formally or informally in a 
liaison capacity, with most of whom constant and 
almost daily contact was maintained: In the Navy 
Bureau of Ordnance, among others, Captain G. L. 
Schuyler, Captain J. S. Champlin, Commander E. A. 
Junghans, Mr. C. B. Green, and especially Lieu¬ 
tenant Commander B. D. Mills, Jr.; in the Army Ord¬ 
nance Department Major H. A. Ellison, Mr. B. E. 
Anderson, Mr. M. C. Miller, Mr. G. E. Stetson, and 
especially Mr. S. Feltman; at Aberdeen Proving 
Ground Mr. R. H. Kent; and at Watertown Arsenal, 
Dr. P. R. Kosting. Above all, without the wise coun¬ 
sel and unfailing patience of Colonel S. B. Ritchie, 
progress would have been much more difficult. 

Special credit is due Dr. J. S. Burlew, the Editor of 
this volume. His skill in arranging for the orderly 
presentation of the joint efforts of a large group of 
authors is matched only by his energy and enthusi¬ 
asm in carrying the undertaking through to the end. 

L. H. Adams 
Chief, Division 1 


vii 


CONFIDENTIAL 

















































































































PREFACE 


O ne of the significant features of the investi¬ 
gation by Division 1, NDRC, of the mechanism 
of gun erosion and of means of attaining hyperveloc¬ 
ity was the close integration of its various phases. 
Not only were the several major projects made up of 
interrelated subprojects, but also there were many 
aspects that were common to two or more of the 
major projects. This situation has been reflected in 
the Division’s Summary Technical Report. 

The report was planned as a logical exposition of 
the different aspects of the main problem denoted by 
its title “Hypervelocity Guns and the Control of Gun 
Erosion.” In relating the final status of the investiga¬ 
tion, no attention has been paid to the way in which 
the program had been organized by projects or as¬ 
signed to different contractors, for such divisions 
were useful only during the administration of the 
program. Instead, chapters have been allotted to 
those subjects considered to be of principal impor¬ 
tance. Thus, the report is a unified one rather than 
merely an assemblage of separate documents. 

At the same time, it is realized that the Summary 
Technical Report is intended to be consulted as a 
reference work, as well as to be read in its entirety. 
Hence an attempt has been made to have each chap¬ 
ter nearly an independent unit. For this reason there 
is some duplication of material, since occasionally it 
has been desirable to consider the same subject from 
more than one point of view. Frequent cross refer¬ 
ences to other parts of the volume have kept such 
duplication to a minimum. 

The entire report was planned in advance and the 
topics to be included in each chapter were specified. 
The order of treatment of these topics and the em¬ 
phasis to be placed on each, however, usually were 
left to the authors. When a topic assigned to one 
chapter was very closely related to the subject mat¬ 
ter of another, the more appropriate place was not 
always apparent until both chapters had been writ¬ 
ten. In some cases the editor, with the authors’ per¬ 
mission, moved material from one chapter to another. 
One feature that is common to many chapters is that 
the first section not only serves to introduce a new 
subject but also furnishes a resume of the chapter. 

The chief source material for the Division 1 STR 
consists of the 143 formal reports, listed in the first 
part of the Bibliography, which were prepared for 
issuance in mimeographed format for the Division by 


the Technical Reports Section of NDRC. These 
classified reports, in which is presented the essential 
information gained by Division 1 contractors, have 
been deposited in a number of Service libraries and 
copies of them are available as microfilms. Because 
of lack of funds and personnel during the final stage 
of the demobilization of OSRD, it was not possible to 
have all these reports mimeographed. Consult the 
paragraph at the beginning of the Bibliography 
concerning their availability. In some instances sup¬ 
plementary details, such as dimensioned drawings or 
additional experimental data, may be found in the 
monthly progress reports or the interim reports from 
contractors. If so, appropriate reference is made in 
the formal report concerning the location of the com¬ 
plete files of these reports. 

Every important topic covered by the Division 1 
formal reports is mentioned in the STR (with a refer¬ 
ence to the appropriate report), so that it is a key to 
the entire work of Division 1 and its contractors. Be¬ 
cause of the close dependence of the STR on this 
body of Division 1 reports, it has not seemed neces¬ 
sary to use quotation marks in those cases where 
sentences have been lifted bodily or whole sections 
have been paraphrased, especially when the author 
of a chapter of the STR was also author of the report 
on which it was based. Even when this is not the 
case, it is hoped that citation of the original report 
will suffice to give proper credit to its author. 

The other sources for the STR are also listed in the 
Bibliography under appropriate headings that indi¬ 
cate the organization issuing the report. All these 
reports are referred to specifically in the text of 
STR, except for British reports' 1 that summarize the 
work of Division 1. 

The principal objective sought in the preparation 
of the STR has been the integration of the knowledge 
gained by Division 1 in its broad investigation of 
erosion and hypervelocity, in order to orient the 
reader with respect to the work described in the 
formal reports from the Division. Chapter 16 on ero¬ 
sion-resistant materials is an especially striking 
example of this feature. It presents results from 
many sources and shows how all of them fit into a 
single program of investigation. Furthermore, the 
preparation of the STR has given opportunity for 

a These reports are listed as the following items in the 
Bibliography: 356, 360, 400, 407, 408, 409, 410. 


CONFIDENTIAL 


IX 



X 


PREFACE 


some of the investigators to review and re-evaluate 
the results of their work as a whole, whereas pre¬ 
viously it had been reported piecemeal. 

In keeping with the general purpose of the Sum¬ 
mary Technical Reports of NDRC, an effort has 
been made to have as many as possible of the chap¬ 
ters present the “state of the art” with respect to the 
subject treated therein. There is considerable vari¬ 
ability in this respect. Thus, at one extreme, Chapter 
5, which deals with the heating of guns during firing, 
is based on Army, Navy, and British reports to as 
great an extent as it is on Division 1 reports and there¬ 
fore approaches a summary of the state of the art. 
Although Chapters 22, 23, and 24 are limited to a 
recital of Division l’s experience, they present the 
state of the art with respect to the development of 
improved machine gun barrels, for the Division has 
been the only contributor in the field. On the other 
hand, a number of the chapters, notably Chapters 4, 
8, 9, 26, 27, and 28, deal with only that limited aspect 
of a general subject that was applicable to Division 
Ts program. 

Because so many of the mathematical equations 
have been taken from reports already issued, a con¬ 
siderable diversity in notation has been introduced. 
It was considered not worth while to make the equa¬ 
tions uniform, for after an equation had been modi¬ 
fied, a reader might find it difficult to compare it with 
related ones in an original report. Therefore a uni¬ 
form system of notation has not been adopted. 

Many of the illustrations used in the STR have 
been taken from previously issued reports of Divi¬ 
sion 1 as indicated by the bibliographical reference 
numbers included in the legends. A few of the illus¬ 
trations have been taken from reports issued by the 
Office of the Chief of Ordnance, to which apprecia¬ 
tion is expressed for permission to reproduce them. 
For some chapters special illustrations were prepared 
by their authors. A number of other figures were 
executed by Mr. Ernest Albert of the H. L. Yoh 
Company, Philadelphia, Pennsylvania. 

The editor wishes to record the fine spirit of coop¬ 
eration displayed by the 25 other authors who con¬ 
tributed to the Division 1 STR. These authors, 
whose names are listed in the Contents of this volume, 
either had taken part in some of the investigations or 
else were members of the Division staff. One of the 
greatest values of the book is the careful appraisal 
they have made of the vast quantity of experimental 
data accumulated by Division 1 contractors during 
its 4-year period of investigation. To each of these 


authors the editor is indebted personally for the 
privilege of having worked with them. 

In addition to having been reviewed by the Divi¬ 
sion Chief, nearly every chapter has been reviewed 
by at least one other person who has an intimate 
knowledge of the subject matter, either through hav¬ 
ing been associated with the investigation itself or 
through having followed the work closely. The 
valuable suggestions made by these reviewers are 
hereby acknowledged. The persons who served Di¬ 
vision 1 in this way and the chapters that they re¬ 
viewed are included in the following list. 

CHAPTER 

W. S. Benedict, National Bureau of Standards. .5, 11 
F. A. Biberstein, The Catholic University of 


America. 27 

H. L. Black, Technical Aide, Division 1, 

NDRC. 6 

W. Blum, National Bureau of Standards.23, 25 

P. H. Brace, Westinghouse Research 

Laboratories. 18 

H. C. Cross, War Metallurgy Committee. 16 

C. L. Faust, Battelle Memorial Institute. 20 

L. H. Germer, Bell Telephone Laboratories.... 21 

J. W. Greig, Geophysical Laboratory, 

C. I. W.23, 24, 29 

K. F. Herzfeld, The Catholic University of 

America. 7 

J. O. Hirschfelder, (formerly) Geophysical 

Laboratory, C. I. W.2, 3, 6 

F. II. Horn, The Johns Hopkins University. . . 14 

W. F. Jackson, E. I. du Pont de Nemours 
and Company.4, 14 

H. S. Jerabek, (formerly) Geophysical 

Laboratory, C. I. W.12, 16, 19 

R. H. Kent, Aberdeen Proving Ground. 8 

J. F. Kincaid, (formerly) Explosives Research 
Laboratory (Division 8, NDRC). 15 

Z. Kopal, Massachusetts Institute of 

Technology. 8 

V. Lamb, National Bureau of Standards.23, 25 

J. J. Lander, Bell Telephone Laboratories. 12 


CONFIDENTIAL 



















PREFACE 


xi 


CHAPTER 

J. P. Magos, Crane Company.22, 24 

.J. W. Marden, Westinghouse Lamp Division. . 18 

H. E. Merwin, Geophysical Laboratory, 

C. I. W.10, 12, 13 

L. W. Nordheim, Consultant, Division 1, 

NDRC. 5 

E. F. Osborn, (formerly) Geophysical 

Laboratory, C. I. W.24, 25 

B. B. Owen, Yale University. 21 

R. M. Parke, Climax Molybdenum Company.. 17 

J. F. Schairer, Special Assistant, Division 1, 

NDRC.17, 18, 20, 21, 26 

F. R. Simpson. 11 

N. H. Smith, The Franklin Institute.15, 28 

W. E. Story, Liaison Office, OSRD. 32 


CHAPTER 


W. F. G. Swann, Bartol Foundation, The 

Franklin Institute.'. 9 

J. A. TenBrook, Special Assistant, 

Division 1, NDRC. 28 

E. G. Townes, Editor, Technical Reports 
Section, NDRC. 8 

W. D. Urry, Geophysical Laboratory, C. I. W. 11 

W. A. Wissler, Union Carbide and Carbon 

Research Laboratories. 19 

N. H. Ziegler, Crane Company. 19 


E. G. Zies, Geophysical Laboratory, C. I. W. 10, 13 

Finally, it is a great personal pleasure to acknowl¬ 
edge the devoted services rendered by both Miss 
Charlotte A. Marsh and Miss Helen M. Watson in 
helping to edit this volume. 

John S. Burlew 
Editor 


CONFIDENTIAL 





















CONTENTS 


CHAPTER PAGE 

Summary. 1 

PART I 

INTRODUCTION 5 

1 The Problem of Hypervelocity by L. H. Adams . 7 

PART II 

BALLISTICS 19 

2 Properties of Powder Gas by F. C. Kracek . . 21 

3 Interior Ballistic Calculations by William S. Bene¬ 
dict .54 

4 Instrumentation for Experimental Ballistic Firings 

by H. B. Brooks .76 

5 Heating of Guns During Firing by H. L. Black and 

G. Comenetz .98 

6 Bore Friction by William S. Benedict . 129 

7 Band Pressure and Related Stresses by H. L. 

Black .152 

8 Exterior Ballistics of Hypervelocity projectiles by 

A. H. Stone .163 

9 Terminal Ballistics of Hypervelocity Projectiles by 

H. S. Roberts and Walker Bleakney .180 

PART III 

GUN EROSION 193 

10 Description of Eroded Gun Bores by Lloyd E. 

Line, Jr .195 

11 Laboratory Methods of Studying Gun Erosion by 

Lloyd E. Line, Jr .219 

12 The Products of Gun Erosion by E. G. Zies and 

C. A. Marsh .244 

13 The Causes of Gun Erosion by C. A. Marsh and 

J. N. Hobstetter ..260 

PART IV 

EROSIVE ACTION OF PROPELLANTS 281 

14 Effects of Constituents of the Powder Gases on 

Gun Steel by W. D. Urry .283 

15 Erosion of Gun Steel by Different Propellants by 

J. N. Hobstetter .308 

PART V 

EROSION RESISTANT MATERIALS 329 

16 Selection of Erosion Resistant Materials for Gun 

Bores by J. F. Schairer .331 


xiii 


CONFIDENTIAL 














XIV 


CONTENTS 


CHAPTER PAGE 

17 Chromium and Chromium-Base Alloys by Helen 

M. Watson .356 

18 Molybdenum by F. Palmer .370 

19 Stellites and other Cobalt Alloys by J. F. Schairer 391 

20 Electroplating by William Blum .408 

21 Vapor-Phase Plating of Molybdenum, Tungsten, 

and Chromium by C. A. Marsh .419 

PART VI 

IMPROVED MACHINE GUN BARRELS 441 

22 Stellite-Lined Machine Gun Barrels by J. F. 

Schairer .443 

23 Nitrided and Chromium-Plated Machine Gun Bar¬ 
rels by E. F. Osborn .458 

24 Barrels both Stellite-Lined and Chromium-Plated 

by J . F. Schairer .473 

25 Pilot Plants for Chromium-Plating Caliber .50 Bar¬ 
rels by V. Wichum and C. A. Marsh .... 485 

PART VII 

HYPERVELOCITY GUNS AND PROJECTILES 501 

26 Short Liners and other Design Features of Gun 

Tubes by William H. Shallenberger .503 

27 Design Features of Projectiles by F. R. Simpson 

and H. L. Black .518 

28 Automatic Gun Mechanism by William H. Shal¬ 
lenberger .538 

29 Sabot-Projectiles by J. S. Burlew .557 

30 Tapered-Bore Guns and Skirted Projectiles by 

Edwin L. Rose .569 

31 Pre-Engraved Projectile with Chromium-Plated 

Bore by Nicol H. Smith .591 

32 The Fisa Protector and Chromium-Plated Bore by 

Nicol H. Smith .609 

33 Practical Hypervelocity Guns by L. H. Adams, 

J. S. Burlew , and E. L. Rose .615 

Glossary.631 

Bibliography.633 

OSRD Appointees.652 

Contracts.653 

Service Projects.656 

Index.657 


CONFIDENTIAL 



















SUMMARY 


HYPERVELOCITY PROBLEM 

D uring recent decades there has been a tend¬ 
ency to use more guns of high power. There has 
not been, however, a corresponding increase in the 
maximum muzzle velocity in common use. This 
maximum has remained at about 3,000 fps, for if a 
higher velocity is used, the rate of erosion is excessive 
and gun life is correspondingly short. 

When, in the summer of 1941, the National De¬ 
fense Research Committee undertook to investigate 
means of making higher velocity guns practical, it 
was recognized that one of the principal hurdles 
would be the control of gun erosion. At the same 
time it was deemed desirable to investigate other 
possible means of attaining this end simply by 
circumventing erosion by a change in design of the 
gun or projectile. Division 1, NDRC, followed con¬ 
currently these two courses of action, accompanying 
them by such investigations of interior ballistics as 
were needed to supplement existing knowledge. 


EROSION STUDIES 

The novel approach that was adopted for the study 
of gun erosion was, in effect, to learn the history of 
the bore surface of different guns that had been fired 
under various conditions. Considerable information 
was obtained from examination of the erosion prod¬ 
ucts found on the bore surfaces of guns worn out in 
service. Of equal importance was the knowledge 
gained from special experiments in which a new gun 
was fired a relatively few rounds and then examined, 
so that the course of the erosion process could be 
followed. A special caliber .50 hypervelocity erosion¬ 
testing gun was developed to make possible many 
routine firings under a variety of conditions. One 
important series of such experiments involved a de¬ 
termination of the relative erosiveness of all the 
common propellants. 

Another group of erosion experiments was under¬ 
taken to observe the effect on a steel bore surface of 
the individual constituents of the powder gases. By 
use of the techniques of physical chemistry it was 
demonstrated that several of the principal constit¬ 
uents react readily with steel at the temperatures 
and pressures prevailing in guns. Mixtures of the 


products resulting from these reactions have a lower 
fusion range than gun steel itself. 

The evidence from these two types of erosion ex¬ 
periments has led to the conclusion that the erosion 
of a steel gun tube results from the reaction of a thin 
layer of steel at the bore surface with the powder 
gases, followed by removal of the reaction products. 
The bore-surface temperature plays a predominant 
role in determining what type of product is formed, 
although the chemical nature of the propellant also 
is a factor. Furthermore, the higher the bore-surface 
temperature the greater the extent to which the al¬ 
tered bore-surface material is softened and even fused, 
and hence more readily removed both by the powder 
gases and by the projectile. In the case of a hyper¬ 
velocity gun fired with a very hot propellant, fusion 
of the gun steel itself occurs, and the molten material 
is blown out of the bore by the powder gases. 

EROSION-RESISTANT MATERIALS 

Early in the course of the erosion studies some ink¬ 
ling of the process delineated in the previous para¬ 
graph was obtained. Thereupon it was tentatively 
concluded as a working hypothesis (which subse¬ 
quent experience confirmed) that severe erosion is in¬ 
evitable in a steel gun tube, regardless of the type of 
steel, when fired under hypervelocity conditions 
with present-day propellants. Thereupon a search 
was begun for nonferrous materials that might be 
erosion resistant. It was soon found that resistance to 
erosion by the powder gases is a property of only a 
few pure metals—in particular, chromium, molyb¬ 
denum, tungsten, and tantalum—and of certain of 
their alloys. It was also learned that a very intensive 
effort would be required to prepare any one of these 
metals in a form suitable for use as a bore-surface 
material. 

Two of the erosion-resistant materials ultimately 
developed by Division 1 were applied to caliber .50 
aircraft machine gun barrels that were used in com¬ 
bat in the latter months of the war in the Pacific. The 
one was stellite, applied as a short breech liner, 
which increased the velocity life of the barrel be¬ 
cause of its high temperature strength and superior 
erosion resistance. The other was chromium electro¬ 
plate applied to the hardened bore surface of the 
barrel, which increased the accuracy life in long 


CONFIDENTIAL 


1 


2 


SUMMARY 


bursts because it was applied in such a way that the 
muzzle end of the bore was “choked” and therefore 
did not become oversize so quickly from the heat of 
firing. After improved barrels of each of these types 
had been in use for a short time, a “combination” 
barrel having choked-muzzle chromium plate ahead 
of a stellite breech liner was developed. This barrel 
had a life at least tenfold greater than that of a steel 
barrel when fired in long bursts (as, for example, in 
strafing) and about a threefold greater life when 
fired in very short bursts. 

A stellite liner was also applied to the Ordnance 
Department’s experimental caliber .60 machine gun 
barrel which has a muzzle velocity of 3,500 fps. The 
use of such a liner may be expected to be mandatory 
for all future machine gun barrels until a superior 
bore surface material is developed. One possible su¬ 
perior material is a chromium-base alloy which can be 
fired with double-base powder, whereas a stellite liner 
melts under these conditions. This alloy requires fur¬ 
ther development to improve its ductility. 

The most promising material so far developed for 
the bore surface of a hypervelocity medium-caliber 
gun is hardened molybdenum. The problem has been 
to learn how to fabricate pieces of ductile molyb¬ 
denum large enough for use as gun liners, and how to 
harden the metal sufficiently so that it will withstand 
the swaging action of a projectile during engraving. 
This goal was achieved in liners for the caliber .50 
erosion-testing gun. Plans were then developed for 
equipment with which to apply the same techniques 
to the fabrication of pieces large enough for liners for 
a 3-in. gun. After OSRD’s contracts had been term¬ 
inated, the Bureau of Ordnance, Navy Department 
continued this work. 

BALLISTIC STUDIES 

In the study of the fundamental causes of gun ero¬ 
sion, knowledge of the physical conditions existing in 
the gun bore during firing was important. Hence the 
process of the burning of powder was studied, a 
means was developed for measuring the temperature 
of the powder gases, and the chemical composition of 
these gases was determined. 

Two other series of measurements dealt with the 
heat input to the bore surface of a gun during firing 
and with the rise of temperature of the bore w^alls, 
especially of a machine gun barrel during the firing of 
a long burst. From these experimental data it was 
possible to arrive at much closer approximations to 


the bore-surface temperature than had been hereto¬ 
fore possible. These results were of inestimable value 
in the formulation of the final explanation of the 
process of erosion. 

Theoretical determinations of bore-surface temper¬ 
atures required the development of an improved sys¬ 
tem of interior ballistic calculations. After the ground¬ 
work in the development of this system had been laid, 
its usefulness for other purposes was realized and ad¬ 
ditional work was done to simplify the computations. 
One advantage of this system over others commonly 
used is that various empirical factors can be allowed 
for explicity. As additional information is gained, it 
can easily be incorporated in the system with a 
resulting improvement in accuracy. This system 
proved to be useful not only in routine ballistic cal¬ 
culations but also in analyses of the ballistic factors 
that enter into a consideration of an increase in 
muzzle velocity. 

Experimental ballistic firings were conducted in an 
effort to obtain more precise information about the 
effect on the behavior of a gun of variables such as 
the weight and composition of the powder charge and 
the projectile band diameter. A new device, termed 
the “microwave interferometer” was developed for 
measurement of projectile movement down the gun 
bore. 

Strain-gauge measurements made on the exterior 
of the gun barrel during these experimental firings 
were correlated with static measurements from a 
study of band pressure. The results were applied to 
both projectile design and gun design. 

Exterior ballistics and terminal ballistics of hyper¬ 
velocity projectiles were subjects of corollary inter¬ 
est but were not investigated extensively by Divi¬ 
sion 1. The theory of the motion of a cone moving at 
high velocity was extended to take into account the 
effect of yaw. Observations were made of the disrup¬ 
tive effect of a hypervelocity projectile upon entering 
a liquid. 

SUBCALIBER PROJECTILES 

The muzzle velocity of any existing gun can be in¬ 
creased considerably by firing from it a subcaliber 
projectile. It seems to be a fairly safe assumption, 
although not actually demonstrated by experiment, 
that the rate of erosion of such a gun is considerably 
less than the rate of erosion of a conventional gun 
firing a standard projectile at the same velocity. 
Hence the use of a subcaliber projectile is a means of 


CONFIDENTIAL 



SUMMARY 


3 


attaining hypervelocity without a marked increase 
in erosion. 

For this reason Division 1 sponsored the develop¬ 
ment of a sabot-projectile, which is one type of sub¬ 
caliber projectile. Considerable attention was paid to 
the use of plastics in such projectiles for the purpose 
of reducing weight; but dimensional instability of 
plastics when exposed to variations in atmospheric 
conditions made it impossible to develop a satisfac¬ 
tory design. Preliminary tests of the final design of an 
all-steel 90-mm sabot-projectile with an 8-lb tungsten 
carbide core, which were made just before the work 
was terminated, indicated that by the use of such a 
projectile the muzzle velocity of the standard gun 
was raised from 2,700 to 3,700 fps. The projectile’s 
accuracy was as good as that of the standard 90-mm 
projectile. 

Another type of subcaliber projectile is a skirted 
one, to be fired from a tapered-bore gun. Experiments 
with tapered-bore guns of different proportions 
showed that the most successful type is one having a 
short unrifled tapered section at the muzzle end. The 
simplest way to provide such a tapered section is by 
attaching a short muzzle adaptor to an existing gun. A 
satisfactory means of doing this for a 57-mm gun was 
developed, together with a design of an armor-pierc¬ 
ing tungsten carbide-cored projectile which had a 
muzzle velocity of 4,200 fps. In this case, also, pre¬ 
liminary tests showed that the accuracy was at least 
as great if not greater than that of the standard round 
fired at a velocity of 2,800 fps from the same gun 
without a muzzle adaptor. 

LINERS AND OTHER FEATURES 
OF GUN DESIGN 

The development of erosion-resistant materials 
naturally led to efforts to utilize them to best advan¬ 
tage. With respect to machine-gun barrels, it was 
demonstrated that the utilization of the full potenti¬ 
alities of stellite liners and of chromium electroplates 
demanded a barrel having a slightly greater wall 
thickness and made of steel having greater strength 
at high temperatures. In preliminary tests such an 
improved barrel showed a severalfold increase in life 
compared with an assembly consisting of the ordi¬ 
nary barrel with the same liner. 

Experience in the application of short breech liners 
to medium-caliber guns was obtained by the develop¬ 
ment of a successful design of replaceable steel liner 
for a 90-mm gun. Such a liner was considered origi¬ 


nally as an emergency means of prolonging the life of 
a steel barrel subjected to severe firing conditions. 
Later it was demonstrated by trial of a 37-mm stel¬ 
lite liner that some features of this design were useful 
as a means of applying an erosion-resistant material 
to a gun bore. 

Although chromium is highly erosion resistant, it 
is difficult to keep an electroplate of chromium on a 
steel gun bore surface. One of the major causes of its 
removal is the high stress concentration set up during 
engraving of the projectile. Two means of preventing 
this action were tried in an effort to enhance the use¬ 
fulness of chromium plate. One was the “Fisa pro¬ 
tector,” which is a thin steel sleeve attached to the 
neck of the cartridge case and extending forward over 
the projectile as far as the bourrelet. It covers the ori¬ 
gin of rifling during engraving and then is extracted 
with the cartridge case. This device has not yet been 
perfected. 

The second means of avoiding undue stress on a 
chromium-plated bore surface is a pre-engraved pro¬ 
jectile. Its successful utilization requires a conven¬ 
ient way of orienting the projectiles prior to chamber¬ 
ing. One self-orienting arrangement that made use of 
pointed lands in the gun was tried in an experimental 
hypervelocity 37-mm gun (velocity: 3,500 fps). The 
severe erosion that occurred at the origin of rifling in 
this particular gun tube might be prevented by the 
use of a thicker chromium plate. 

A gun design project that was separate from the 
rest of the Division’s activities was the development 
of an improved mechanism for a 20-mm automatic 
cannon. Although a completely satisfactory design 
had not been developed by the time the project was 
terminated, firing tests of the latest model indicated 
the hope that one might be achieved eventually. 

EXPERIMENTAL HYPERVELOCITY GUN 

After Division l’s investigations had yielded a ra¬ 
tional explanation of the phenomenon of gun erosion 
and had pointed the way toward the development of 
erosion resistant materials, an attempt was made to 
apply this new knowledge to the development of a 
practical hypervelocity gun. The first step in this di¬ 
rection was to have been an experimental 90-mm gun 
firing at a muzzle velocity of at least 4,000 fps. Two 
tubes for this gun, chambered and bored but not 
rifled, were prepared before the termination of the 
Division’s activities and then were turned over to the 
Army Ordnance Department for further develop- 


CONFIDENTIAL 



4 


SUMMARY 


ment. According to the original plans these tubes 
were to have been chromium plated and were to have 
fired pre-engraved projectiles, and thus make use of 
the experience gained in the trial of the 37-mm tubes. 

The next step in this program was to have been the 
insertion of a molybdenum liner in another tube hav¬ 
ing the same ballistic characteristics. As a prelimin¬ 
ary, plans were made for the fabrication of a molyb¬ 
denum liner for the Navy’s new 3-inch/70-caliber gun. 
Upon cessation of Division l’s sponsorship, this proj¬ 
ect was taken over by the Navy Bureau of Ordnance. 

Means are thus at hand for an immediate study of 


some of the characteristics of medium-caliber hyper¬ 
velocity guns. What is needed next is a careful evalu¬ 
ation of the possible ways of utilizing erosion-resist- 
ant materials in such guns by actual trial of them. 
Firings should be conducted to determine the opti¬ 
mum ballistic conditions, particularly means of in¬ 
creasing the density of loading. Finally, then, by 
combining such information with that concerning the 
effects on different targets of hypervelocity projec¬ 
tiles, it should be relatively simple to decide on a de¬ 
sign of gun best adapted to serve any particular 
Service requirement. 


CONFIDENTIAL 



PART I 

INTRODUCTION 



P 



On what strange stuff ambition feeds! 

—Eliza Cook 
“Thomas Hood” 


CONFIDENTIAL 
















Chapter 1 

THE PROBLEM OF HYPERVELOCITY 

By L. H. Adams a 


I t is the purpose of this chapter to indicate briefly 
the nature of the hypervelocity problem, to sum¬ 
marize the advantages of and the difficulties with 
higher muzzle velocities in guns, to indicate the meth¬ 
ods by which some of the difficulties have been ob¬ 
viated, and to outline the various means that have 
been employed to make guns of greatly increased per¬ 
formance practicable. Throughout the volume as a 
whole these same broad topics form the basis of de¬ 
tailed presentation. The present chapter concludes 
with some remarks comparing the status of the hyper¬ 
velocity problem in 1946 with that in 1941. 

i-i NATURE OF THE PROBLEM 

1,1,1 Some Elementary Considerations 

A gun is a device for converting the chemical energy 
of a powder charge into mechanical energy of a mov¬ 
ing projectile and for launching the projectile toward 
a target in an accurately specified direction. Thus, it 
serves two principal purposes. As an engine, its me¬ 
chanical efficiency compares favorably with the best 
turbines and internal combustion engines. For exam¬ 
ple, a representative gun with a 3-in. bore and a muz¬ 
zle velocity of about 2,700 fps converts 29 per cent of 
the powder energy into kinetic energy of the projec¬ 
tile. (Compare Table 7, Chapter 6.) Of the remaining 
71 per cent, about 9 per cent is accounted for by heat 
conducted into the gun barrel, the remainder being 
accounted for mainly by the heat and kinetic energy 
of the hot gases ejected from the gun. 

Under favorable conditions, a greater proportion of 
the powder energy ma}^ be converted into energy of 
the projectile; but in a gun operating at extreme muzzle 
velocities, the mechanical efficiency is lower, mainly 
because more energy is lost in the effluent gases. For¬ 
tunately, this falling off in efficiency does not become 
important until velocities very much higher than 
conventional ones are reached. For example, the cal¬ 
culated value for a 90-mm gun operating at 4,250 fps 
is only about 10 per cent lower than that for an exist¬ 
ing gun of the same caliber operating at 2,700 fps. At 

a Chief, Division 1, NDRC. (Present address: Geophysical 
Laboratory, Carnegie Institution of Washington.) 


still higher velocities, the mechanical efficiency of a 
gun begins to drop more rapidly, and at 6,000 fps 
muzzle velocity would not exceed two-thirds of that 
of existing guns, the velocity of which seldom exceeds 
half this magnitude. 

A variety of complex factors, discussed in Chapters 
2, 3, 5, and 6, determine the overall efficiency and 
hence the energy and muzzle velocity of a given pro¬ 
jectile when shot from a given gun with a given pow¬ 
der charge. Without going into details, it suffices for 
the moment to note that within the range of custom¬ 
ary designs and for a specified powder, the muzzle 
velocity depends primarily on the ratio of powder 
weight to projectile weight, and may be varied within 
wide limits simply by changing this ratio. That is, 
with a given gun, higher velocities may be obtained 
either by providing for more powder capacity or by 
decreasing the weight of the projectile. Contrary to 
what has sometimes been assumed, the limiting velo¬ 
city of a projectile is not the velocity of sound in the 
powder gas at the temperature and pressure existing 
in the gun during firing; velocities of 9,000 fps (about 
three times the pertinent sound velocity) have been 
obtained experimentally. (See Section 3.5.2.) Rather, 
the limit is a practical one, imposed by other consid¬ 
erations which are listed and discussed briefly in 
Section 1.3. 

In order that the gun may fulfill satisfactorily its 
second principal purpose, namely, that of a projectile 
launcher, it is necessary that the projectile be ejected 
with a predetermined velocity and direction. For pro¬ 
jectiles stabilized in flight by spinning, there is the 
added requirement that the rate of spin be adequate 
for stabilization. That modern guns have a remark¬ 
able accuracy may be emphasized by comparing their 
precision with that of the human eye. The best guns, 
under favorable conditions, show a mean dispersion 
at the target of less than one-half mil, or about 1 min 
of arc, which happens to be the same as the limit of 
resolution of the eye under conditions of good light¬ 
ing and normal vision. Thus, with a good gun, the 
gunner should be able to see only with difficulty the 
separate hits on the target at any range. 

When a gun is fired, the bore surface deteriorates, 
and the consequent alteration of the above-mentioned 


CONFIDENTIAL 


7 



8 


THE PROBLEM OF HYPERVELOCITY 


ballistic factors gradually impairs the accuracy and 
finally brings the gun to the end of its useful life. It 
was stated in the preceding paragraph that, viewed 
as an engine, a gun is an efficient mechanism. But, 
whereas the operating life of a gasoline engine is meas¬ 
ured in thousands of hours, that of a gun is measured 
in minutes or even seconds. This follows from the cir¬ 
cumstance that a gun is actually operating only while 
the projectile is passing down the bore. In a 12-in. gun 
of the usual design the travel time in the bore is about 
30 msec; and if we multiply this time interval by 350 
(a reasonable estimate for the useful life of a 12-in. 
gun) we obtain approximately 10 sec as the operating 
life. It is a curious fact that the corresponding figure 
for guns of various sizes is of the same order of mag¬ 
nitude, provided that the muzzle velocity remains 
approximately constant and the gun is fired slowly; 
for with smaller guns the travel time of the projectile 
down the bore is shorter, and (ordinarily) the usable 
number of rounds is correspondingly greater. It ap¬ 
pears, therefore, that a gun is an efficient thermody¬ 
namic machine but that its life, whether measured in 
number of rounds or time of operation, may be dis- 
couragingly small. All efforts to make guns more 
effective must take into account the element of useful 
life. 

1,1,2 Trend of Muzzle Velocities 

in the Past 

Cannon were in use as early as the 14th century, 
at which time stone shot were fired from wrought iron 
barrels, the propellant being a crude form of black 
powder. Little information is available concerning 
the performance of these weapons, but it is unlikely 
that the muzzle velocities were more than a few hun¬ 
dred fps. Beginning in 1520, cast iron round shot be¬ 
gan to be used and muzzle velocities were gradually 
increased. Records of Italian guns of 1746 show muz¬ 
zle velocities in excess of 1,600 fps. A notable advance 
in ordnance began to take place about 1850 with the 
discovery and application of progressive-burning pow¬ 
ders. Somewhat later, black powder, which had held 
the field in artillery for over 500 years, was displaced 
by nitrocellulose powder. During this same period, 
elongated rifled projectiles were developed, and by 
the year 1900 several nations were employing large¬ 
sized guns with muzzle velocities of 3,000 fps. Sub¬ 
sequent to that year, there has been little or no 
change, except perhaps the use of a greater propor¬ 
tion of high-velocity guns. 


It may be instructive to note the rate of change of 
muzzle velocities for cannon throughout the past. For 
the 500 years ending in 1850 the average increase was 
about 5 fps per year, while for the 50 years following 
that date the average increase was of the order of 30 
fps per year. During the 4 decades subsequent to 
1900, the change was essentially zero. 

In this connection, it is pertinent to take note also 
of the increase in rate of fire that has accompanied 
the increase in velocity. 510 The primitive guns were 
muzzle loading and required many minutes to reload 
and fire. With the advent of breech loading, and later 
of cased ammunition, the time per round was signif¬ 
icantly reduced; and in the fairly recent past, auto¬ 
matic mechanisms have greatly increased the rate of 
fire. We now have 5-in. cannon firing at 20 rounds per 
minute and small arms (machine guns) firing at rates 
of 1,200 rounds per minute. Contrasted with muzzle 
velocity, rate of fire does not seem to have reached a 
corresponding state of stagnation. As mentioned in 
Section 1.3.2, these two aspects of improved gun per¬ 
formance involve features that are closely interrelated. 

1.1.3 The Term Hypervelocity 

There has been a tendency to make a rough divi¬ 
sion of artillery into low-velocity and high-velocity 
guns, the former class including velocities not much 
higher than 2,000 fps and the latter running up to 
3,000 fps or a little above. Partly because there seemed 
to be unusual difficulties in the practical attainment 
of guns operating at muzzle velocities much in excess 
of 3,000 fps, it has been deemed convenient to use a 
special designation for guns of very high velocities. 
The term hypervelocity has been used; and in this vol¬ 
ume hypervelocity guns are considered by definition 
to be ones operating at muzzle velocities of 3,500 fps 
or higher. b 

1.1.4 The Problem Defined 

A study of the past almost inevitably arouses our 
curiosity as to whether the long upward trend in 
muzzle velocities has properly come to an end, or 
whether further advances are both possible and prof¬ 
itable. We are confronted with a problem, the basis 
of which may be formulated by a series of questions 
as follows. 

b This usage follows that adopted by the Ordnance Board of 
Great Britain, the limiting velocity having been set at 3,500 
fps by the Hypervelocity Panel at its meeting on May 28, 1942. 


CONFIDENTIAL 




ADVANTAGES OF HIGHER PROJECTILE VELOCITIES 


9 


(1) What are the advantages of hypervelocity 
guns? 

(2) What obstacles are in the way of attaining 
higher muzzle velocities than have heretofore been 
considered practicable? 

(3) What are the physical and chemical factors in 
such limitations? 

(4) By what new knowledge can the difficulties be 
resolved? 

(5) How can the results of such knowledge be 
translated into weapons of increased effectiveness? 

Anticipating at this point some of the conclusions 
mentioned in the following paragraphs and explained 
more fully elsewhere in other chapters, we may note 
that hypervelocity does have worthwhile advantages 
and does involve serious difficulties; and that of the 
latter the outstanding one is erosion of the gun bore. 
The core of the problem is erosion. As implied by the 
title of this volume, if we can find the cause and cure 
of gun erosion and other deteriorations of the bore 
surface—or, if not a cure, a means of obviating the 
effects—we shall have gone a long way toward the 
complete solution of the hypervelocity problem. 

i-2 A PRIORI ADVANTAGES OF 
HIGHER PROJECTILE VELOCITIES 

121 General 

It should be fairly obvious that a projectile at high 
velocity will outperform one at low velocity. Not so 
obvious perhaps is the increased effectiveness, in sev¬ 
eral respects, for the same energy or for the same 
weight of equipment. More exactly, the comparison 
is made between two projectiles of different calibers, 
either having the same muzzle energy or being fired 
from two guns having the same total weight, includ¬ 
ing associated equipment. As a means of summarizing 
the advantages of hypervelocity, it is convenient to 
group them in three categories: greater range; in¬ 
creased armor penetration; and increased probability 
of hitting the target. (Compare Section 33.2.) 

In addition to these principal advantages, there are 
others, less well defined, that may merit careful eval¬ 
uation. For example, as explained in Section 9.1, hy¬ 
pervelocity bullets show a strange disruptive effect 
when impinging on a vessel filled with liquid. For 
each of the aspects of hypervelocity, there is a limit¬ 
ing velocity beyond which, for practical reasons, it is 
not profitable to go, but it is now quite certain that 


the limit is well above the Velocities that have been 
used heretofore. 

122 Greater Range 

In the ideal case of negligible air resistance, the 
trajectory is a simple parabola and the maximum 
range of a projectile varies as the square of the initial 
velocity. 0 This applies to both the horizontal and the 
vertical range. Actually, the relationship is a very 
complex one, owing to the large retarding effect of the 
atmosphere on the projectile; but the range neverthe¬ 
less increases rapidly with increase of initial velocity 
and for a 50 per cent increase in muzzle velocity the 
maximum range of a large gun would be extended by 
something like a ratio of two to one. If a comparison 
is made on the basis of maximum effective range, it is 
usually found that this quantity also increases rapidly 
with muzzle velocity. 

Even a moderate increase of horizontal range might 
be of great value if the guns of an enemy could there¬ 
by be outranged. Moreover, greater vertical range 
would clearly be of great value in coping with high¬ 
flying airplanes. These advantages are of course to be 
assessed against a variety of tactical situations, but it 
is likely that under some conditions the advantages 
would offset a considerable increase in weight and 
bulk of equipment. 

123 Improved Armor Penetration 

Measurements of the penetration of armor plate by 
projectiles of various sizes and types have shown that 
the thickness of plate that can be penetrated increases 
more rapidly than the striking velocity, being propor¬ 
tional to approximately the three-halves power of the 
velocity. This relation is known to hold approximately 
up to velocities somewhat above 3,000 fps, provided 
that the projectile does not shatter upon impact, as 
brought out in Section 9.2. With steel projectiles, 
such shattering without penetration is likely to take 
place at velocities somewhat above the velocities at 

c Constant force of gravity and a plane reference surface 
is assumed also. Under these ideal conditions, a shot at 
an angle of elevation of 45 degrees and an initial velocity of 
3,000 fps would travel to a distance of 54 miles, and a shot 
projected vertically upwards with the same initial velocity 
would reach a height of 27 miles; but the effect of air resistance 
is so great that for a standard 4.7-in. projectile, with a muzzle 
velocity of 3,000 fps, the maximum horizontal range is only 
about 15 miles (attained with an angle of elevation of 47^ 
degrees). The corresponding maximum vertical range is 11 
miles. 


CONFIDENTIAL 




10 


THE PROBLEM OF HYPERVELOCITY 


which they would just perforate the plate. d At still 
higher velocities, the projectiles may penetrate even 
though they shatter, but there may be a range of dis¬ 
tances (and striking velocities) at which a particular 
projectile is ineffective against armor. This might 
appear to limit seriously the usefulness of hyperveloc¬ 
ity guns. But it should be noted that at all except 
short distances from target to gun, the striking 
velocity usually will have diminished to such an ex¬ 
tent that it will be well below the velocity at which 
shattering takes place. 

Further encouragement is afforded by the behavior 
of projectiles having cores of cemented tungsten car¬ 
bide. Such projectiles show greater penetrating power 
than ordinary steel projectiles, as may be seen from 
an Ordnance Department test of a 1.5-in. (37-mm) 
projectile having a tungsten carbide core, which pene¬ 
trated a 3% plate of Class “B” armor at 20 degrees 
incidence at a velocity of 3,440 fps, whereas a pro¬ 
jectile of similar design, but having a steel core, did 
not penetrate the same plate at a striking velocity of 
over 4,000 fps. Above a certain velocity, tungsten 
carbide cores break up when passing through armor 
plate. The older cores disintegrated into fine particles, 
but it is now known that cores can be fabricated in 
such a way that on passing through armor they break 
up into a moderate number of fragments which for 
antitank use are effective in putting the tank out of 
action. 

Results of experiment and calculation indicate that, 
by the use of hypervelocity projectiles, materially in¬ 
creased effectiveness is achieved for the same total 
weight of gun and mount. Put in another way, it is 
readily practicable to obtain a desired effect with about 
one-half the total weight of the usual equipment, and 
thus with a corresponding increase in mobility. 190 

12 4 Increased Probability 

of Hitting Target 

Higher velocities mean shorter times of flight and 
flatter trajectories; both factors are important be¬ 
cause they increase the chance of hitting the target. 
In antitank warfare, 145 a flatter trajectory is advan¬ 
tageous because the resulting decrease in the angle 
of impact reduces the chance of missing the target 
through imperfect knowledge of the range. If an error 

d Recently, steel armor piercing projectiles of improved de¬ 
sign have been shown to penetrate armor 3 calibers in thick¬ 
ness, without shattering, at velocities in the neighborhood of 
4.800 fps at an angle of 30 degrees. (See Section 9.2.1.) 


is made in estimating the distance from gun to target, 
the flatter the trajectory the less will be the vertical 
distance from the point of aim. This may well make 
the difference between hitting an object of a given 
height and missing it altogether. As an example, we 
may note that for a tank 8 ft high and for a conven¬ 
tional 75-mm projectile with a muzzle velocity of 
2,030 fps, the allow able error in estimating the range 
at 1,500 yd is only 100 yd, while with a tungsten car¬ 
bide subcaliber projectile shot from the same gun 
with a muzzle velocity of 3,550 fps, owing to the flat¬ 
ter trajectory, the allowable error at the same range 
is 400 yd. 

A reduction of the “lead” required in aiming at a 
moving target is one consequence of the shorter time 
of flight. For the tw r o projectiles just referred to, the 
respective leads to be applied by the gunner in the 
case of a tank at 1,500 yd moving at a speed of 30 
mph are 105 ft as compared with 60 ft. In the case of 
aircraft gunnery, the reduction of lead is also impor¬ 
tant; but here the advantages of hypervelocity are 
difficult to assess, because of various factors, espe¬ 
cially the limitation in weight of equipment. 146 - 325 - 375 

Short times of flight are particularly desirable in 
firing at fast-moving targets such as airplanes, the 
advantage being most noteworthy when the target 
resorts to dodging tactics. 164 Many mathematical 
analyses have been made of the probability of hitting 
a plane from the ground in relation to the velocity of 
the projectiles. The answer depends upon the assumed 
conditions of the problem. Depending on these condi¬ 
tions, the number of hits for a given number of shots 
is found to be proportional to a power of the velocity 
that varies from the third to the sixth for targets 
moving in three dimensions. Even on the most con¬ 
servative basis, that is, variation with the third 
power, a 50 per cent increase in velocity more than 
triples the number of hits. 

FACTORS LIMITING USABLE MUZZLE 
VELOCITY 

131 Individual Problems to be Solved 

As indicated above, there is no particular difficulty 
in obtaining projectile velocities very much greater 
than those in conventional guns; but there are certain 
practical considerations that seriously limit the veloc¬ 
ities that can be usefully employed. Among these 
factors, the immediately obvious is the w earing away 
of the material at the bore surface, especially when 


CONFIDENTIAL 




FACTORS LIMITING USABLE MUZZLE VELOCITY 


11 


muzzle velocities are much above 3,000 fps. Other 
effects such as the overall heating of the gun barrel, 
the decreased mechanical efficiency of the gun, and 
the loss of projectile velocity in flight, add to the com¬ 
plexities of the hypervelocity problem. Means must 
be found to combat them all by solving separately 
the problems that arise, for the deleterious phenom¬ 
ena are rapidly enhanced whenever an attempt is 
made to achieve higher muzzle velocities. 

At each stage in the history of ordnance some one 
or more of the factors that limit the usefulness of a 
gun has been predominant. As improvements in ma¬ 
terials and in methods of fabrication have led to bet¬ 
ter ways of making guns and their ammunition, it has 
been necessary for the gun designer periodically to re¬ 
evaluate the problem in the light of the tactical needs 
of the moment. On the basis of such analyses com¬ 
promises have been reached in the design of guns for 
particular purposes. Thus shortly before World War 
II design studies were made at Aberdeen Proving 
Ground as to the best means of improving the effective¬ 
ness of antiaircraft guns 550 and of antitank guns. 205,551 
Although these studies indicated that considerable 
advantage would be gained in both applications by 
increasing the muzzle velocity to well above 3,000 fps, 
the new guns that were built later for these purposes 
perforce had lower velocities. Lack of a means of ero¬ 
sion control still prevented the use of otherwise supe¬ 
rior designs. 

132 Erosion e 

During the firing of a gun, the powder burns and is 
converted into white-hot gas under high pressure, 
which propels the projectile down the bore. At the 
same time, the mass of compressed gas mainly follows 
the projectile; and what may be called a condensed, 
swiftly-moving flame heats, chemically alters, and 
scours the bore surface. Some gas may leak past the 
projectile and produce locally intensified effects. In 
addition, the rubbing of the projectile, or its rotating 
band, against the inner wall of the gun tube may 
abrade the surface; and other actions such as abra¬ 
sion by unburned powder grains may take place. 

After a few rounds at high muzzle velocities, or 
after many rounds at lower velocities, enlargement of 
the bore becomes readily evident. The effects are 
more noticeable at the two ends of the barrel, but are 
not confined to these localities. At or near the origin 


e This subject is discussed at length in Part IV. 


of rifling, the damage to the bore surface is a max¬ 
imum. This “origin erosion”—or, as it is sometimes 
called, “breech-end erosion”—may amount in a worn 
gun to several per cent of the bore diameter. Although 
the erosion per round is measured in fractions of a 
thousandth of an inch, or at most, a very few thou¬ 
sandths, the amount of metal lost is sometimes im¬ 
pressive, particularly in large guns. For example, in a 
12-in. gun firing at 2,600 fps, nearly a pound of mate¬ 
rial may be removed from the bore surface during a 
single round. The erosion drops off rapidly in the for¬ 
ward direction and rarely is significant beyond about 
10 calibers distance from the origin. With either sep¬ 
arate-loading or fixed ammunition, a gun in an eroded 
condition produces less muzzle velocity unless the 
powder charge is increased. Another disadvantage re¬ 
sults when the ammunition is of the “fixed” type, for 
with such ammunition, the initial free run of the pro¬ 
jectile in a badly eroded gun may cause stripping of 
the rotating band and damage to the fuze. 

Origin erosion in a given gun is directly related to 
the weight of powder and only indirectly to the muz¬ 
zle velocity, but, in general, larger charges are neces¬ 
sary for higher velocities; and for only a 50 per cent 
increase in muzzle velocities, say from 2,800 fps to 
somewhat above 4,000, it can be expected that origin 
erosion will be increased by several times with a cor¬ 
responding decrease in gun life unless means are 
found to cope with the erosion. 

Near the muzzle, the total erosion is notably differ¬ 
ent in character and magnitude. It is greatest at the 
muzzle where the radial extent is usually only a small 
fraction of the origin erosion in the same gun, and it 
decreases in the rearward direction, usually becoming 
zero at a distance one-third or more of the distance to 
the breech. Although the enlargement of the bore at 
the muzzle is small in comparison with that at the 
breech end of the tube, muzzle erosion (sometimes 
called muzzle wear) nevertheless may lead to trouble¬ 
some dispersion of shots by producing an initial yaw 
of the projectile. It is now known as a result of studies 
undertaken by Division 1 that the rate of muzzle ero¬ 
sion increases exponentially with the muzzle velocity; 
and it is probable that at very high velocities this 
type of erosion rather than the more familiar origin 
erosion constitutes the more serious limitation. 

133 Heating of Gun Barrel f 

By transfer of heat from the powder gases, the 

f See Chapter 5 for further details. 


CONFIDENTIAL 





12 


THE PROBLEM OF HYPERVELOCITY 


temperature of the gun barrel rises to an extent de¬ 
pendent on a variety of conditions. The temperature 
rise is, of course, greatest at the inner surface, where a 
resulting high temperature not only enhances all 
erosive effects but also may soften the inner layers of 
the wall so that the lands are flattened by action of 
the projectile, and the rifling no longer imparts the 
proper spin to the projectile. When successive shots 
are fired at short intervals, there is an accumulation 
of heat so that the barrel wall as a whole may be over¬ 
heated. In machine guns after prolonged bursts even 
at conventional velocities, the temperature of the 
barrel may increase to as much as 800 C; so that the 
barrel may swell and eventually blow up. In rapid- 
fire guns of all calibers, operating at hypervelocities, 
the heating of the gun barrel is one of the major diffi¬ 
culties to be overcome before such guns become en¬ 
tirely practicable. From what has been said, it is evi¬ 
dent that the high rate of fire intensifies the difficul¬ 
ties encountered with hypervelocity; and it is there¬ 
fore not practicable to separate completely these two 
aspects of the broader problem. 

1.3.4 Projectile Stabilization 8 

Ordinarily, projectiles are afforded stability in 
flight by being given a rapid spinning motion which is 
caused by the engagement of helical rifling grooves of 
the gun with a rotating band on the projectile. It can 
be shown both theoretically and experimentally that 
in order to obtain the usual rates of spin (varying 
from several thousand revolutions per minute in large 
guns to over 100,000 rpm in small ones), the stresses 
set up in the banding material by its interaction with 
the rifling often are close to the maximum strength of 
the commonly used materials for rotating bands. 
Moreover, for a given projectile, the requisite rate of 
spin for stability in flight is very nearly proportional 
to the velocity. It follows, therefore, that with con¬ 
ventional projectiles band failure is likely to limit the 
muzzle velocities that can be usefully employed, and 
that the broad hypervelocity problem includes among 
its subsidiary problems that of providing adequate 
stability. 

1.3.5 Other Mechanical Limitations 

In the usual high power guns for military applica¬ 
tion, internal pressures of the order of 50,000 psi are 


* Compare Section 8.3. 


employed. The design of hypervelocity gun tubes 
.would be simplified if still higher pressures could be 
used. (Section 33.3.9) This would require tubes with 
higher strength materials or with thicker walls, or 
both. Higher velocities for a projectile of a specified 
weight require larger powder charges unless a more 
powerful propellant is used. As noted in the preceding 
text, very high muzzle velocities are attained at the 
cost of lessened efficiency at which the energy of the 
powder is converted into kinetic energy of the pro¬ 
jectile, and therefore powder weight and powder- 
chamber capacity increase faster than the muzzle 
energy. A more troublesome limitation is introduced 
by the fact that larger powder charges almost in¬ 
evitably require longer, heavier, and more cumber¬ 
some gun tubes. Clearly, the overall weight of the 
tube and mount is an important factor that enters 
into considerations involving the advantages of hy¬ 
pervelocity and into the design of guns having opti¬ 
mum effectiveness. 

It is evident that when all, or a sufficient number, 
of the individual problems have been solved and the 
limiting factors properly taken account of, there re¬ 
mains a careful balance between conflicting elements 
of design, if a satisfactory use is to be made of hyper¬ 
velocity. 

14 STATUS OF PROBLEM IN 1941 

1,4,1 Divergent Views on Advantages of 
Hypervelocity 

To those scientists and engineers who in 1941 made 
their first acquaintance with ordnance matters, the 
most striking aspect of the situation was the lack of 
agreement among qualified observers as to the ad¬ 
vantages of guns having muzzle velocities much 
higher than those being currently employed. On the 
one hand, there were circumstances 425 that indicated 
pronounced advantages, as brought out in Section 
1.2; while on the other hand, certain lines of reasoning 
indicated that higher velocities would offer little or 
no improvement in overall performance. 146 ’ 325 

The need for increased range, both horizontal and 
vertical, had been urged in certain quarters, and in 
particular it was evident that a higher vertical range 
would be desirable as a protection against airplanes 
which were flying at ever-increasing altitudes. 550 

Prior to 1941, some attention had been given to the 
presumable advantages of machine guns with higher 
muzzle velocities. The need for such guns had not 


CONFIDENTIAL 




APPROACHES TO ATTAINMENT OF HYPERYELOCITY 


13 


been clearly formulated nor had the desirability of 
developing machine guns that would fire long bursts 
without becoming useless been recognized. 

142 Imperfect Understanding 

of Nature of Erosion 

Despite the numerous investigations over a period 
of years on gun erosion, some of which were con¬ 
ducted with great skill and perseverance, there was 
little real understanding of the processes by which the 
bore of a gun is eroded during firing. To be sure, a 
number of suggestions had been put forward as to the 
principal causes of erosion. 15 - 16 Among these were 
abrasion of the surface by unburned powder grains, a 
scouring of the surface by the blast of powder gases, a 
melting of the surface material and subsequent blow¬ 
ing away, chemical action by the constituents of the 
hot gases, friction between the bore and projectile, 
and a variety of other factors. The results of previous 
efforts to reduce erosion had not been encouraging. In 
the belief that materials other than steel were not 
practicable for gun barrels, most of the efforts had 
been devoted to the development of steels of better 
performance. It was thought in those days that an 
improvement of something like 5 per cent in erosion 
resistance would represent a worthwhile goal. 244 

The general attitude prior to the war may be char¬ 
acterized as a tendency to take a somewhat pessi¬ 
mistic view of the erosion problem and to minimize 
the advantages of the higher velocities that would be 
practicable if erosion could be cured or by-passed. 

143 NDRC Attacks Problem in 

Comprehensive Way 

Initially, the National Defense Research Commit¬ 
tee [NDRC] undertook projects of direct and imme¬ 
diate applicability to military needs. Although there 
was an obvious need for guns having better perform¬ 
ance, it was generally known that serious difficulties 
stood in the way of realizing guns with higher muzzle 
velocities, and NDRC quite naturally hesitated to 
spend time, effort, and money on problems of uncer¬ 
tain value for the current emergency or for the im¬ 
pending war. Gun erosion was the key to the situa¬ 
tion, and preliminary consideration of the whole 
matter raised grave doubts as to whether it would be 
proper or profitable for NDRC to attempt to engage 
upon a program for the solution of the age-old erosion 
problem. It is not to be wondered at that the Armed 


Services should have had little hope that this prob¬ 
lem could be satisfactorily solved or should offer only 
faint encouragement for taking up a hypervelocity 
project as a wartime measure. 

The situation eventually crystallized in a decision, 
largely as a result of the views of the British, who 
strongly urged that NDRC undertake broad problems 
in ordnance and placed as number one on their list of 
projects the practical attainment of hypervelocity 
guns and projectiles. In particular, for antiaircraft 
use, they emphasized the need for a gun that would 
fire a projectile having a time of flight of only 3 sec to 
10,000 ft—which would require a muzzle velocity of 
from 4,000 to 5,000 fps. 

In the summer of 1941, NDRC decided to embark 
on a broad program, and for this purpose, Section 
A-A, later Division 1, was created, its assignment be¬ 
ing “gun erosion and other ordnance problems,” a 
title intended to camouflage the real objective, which 
was the development of practical hypervelocity guns 
and in general the improvement of gun performance. 
The program was intended to be a long-range one 
with the initial emphasis on basic research, with the 
understanding that the immediate trial of promising 
devices and even the playing of hunches were not to 
be excluded. This ambitious venture was entered into 
with the full knowledge of the complexities of the sub¬ 
ject, and a realization of the difficulties that even ex¬ 
perienced investigators would encounter, if they un¬ 
dertook to solve problems in a field with which they 
were not familiar and in which ordnance experts had 
made little progress for many years. But it was the 
duty as well as the privilege of NDRC to concern it¬ 
self with ordnance projects of any kind that seemed 
worthwhile, and it was evident that a thorough ex¬ 
ploration of the possibilities of vastly improving the 
existing guns was a necessary “hedge’’ against a long- 
continuing war, the hypervelocity program being in 
the nature of essential insurance against the adverse 
effects of increasing technical proficiency on the part 
of our potential enemies. 

I- 5 APPROACHES TO PRACTICAL 
ATTAINMENT OF HYPERVELOCITY 

151 Classification of Methods 

Higher velocities ordinarily are accompanied by 
greater erosion. The practical attainment of hyper¬ 
velocity therefore involves principally the discovery 


CONFIDENTIAL 



14 


THE PROBLEM OF HYPERVELOCITY 


and application of methods for avoiding or circum¬ 
venting erosion so that the gun may have a reasonably 
long life. Several lines of attack were apparent as the 
program of research and development began to be 
formulated. Erosion may be combatted by an ero¬ 
sion-resistant liner or coating. A suitable protection 
might also be given by a film of some sort applied to 
the bore surface. Inasmuch as the most seriously 
eroded part of the bore is near the origin of rifling, 
this part of the bore surface could be protected from 
the powder gases by a thin sleeve attached to and ex¬ 
tracted with the cartridge case. Lightweight projec¬ 
tiles may be used in a standard gun, a procedure 
which partially solves the erosion problem. Finally, 
propellants that have a lower flame temperature and 
less chemical activity might significantly diminish 
the erosive effect. The specific devices and methods 
that, after preliminary inspection of the problem by 
Division 1, seemed promising for investigation are 
given in the following paragraphs and are described 
in detail in other chapters of this volume. 

152 Replaceable Steel Liners 

It was said somewhat facetiously in the early days 
of Division 1 that the way to solve the hypervelocity 
problem was to let the gun wear out, throw it away, 
and get a new one. More seriously, a liner could be 
discarded; and it was judged worthwhile to investi¬ 
gate thoroughly the practibility of developing a short 
steel liner, which could be inserted in the bore at the 
breech end and, when badly worn, promptly replaced 
by a new one. This development is described in Sec¬ 
tion 26.3. 

15 3 Erosion-Resistant Liners and Coatings 

The selection and development of materials that 
will resist gun erosion obviously require a knowledge 
of the nature of gun erosion and the principles by 
which it is possible to combat it. Parts III and IV of 
the present volume are taken up with the description 
of the basic research applied to this part of the prob¬ 
lem. It was necessary to have better information con¬ 
cerning the behavior of the powder gases in a gun and 
concerning other features of interior ballistics. As a 
result of the intensive studies on erosion, it was pos¬ 
sible at an early stage of the investigation to define 
accurately the physical, chemical, and mechanical 
requirements for usable erosion-resistant substances, 
and to make a preliminary selection of erosion-resist¬ 


ant materials for development and trial. As will be 
shown in Parts V and VI, primary attention was 
focused on chromium, on molybdenum, on alloys con¬ 
taining dominant amounts of these metals, and also 
on high-cobalt alloys. Some of these, as for example 
chromium, seemed more suitable for application as 
electroplated coatings. Others were more promising 
for application as liners. 

154 "Fisa” Protector 

A suggestion that came from abroad formed the 
basis of another approach to the practical attainment 
of hypervelocity. This took the form of a thin sleeve 
to be attached to the forward end of the cartridge 
case enclosing the rear portion of the projectile and 
extracted with the case after firing. This device seemed 
worth developing because the erosion in a gun is most 
noticeable near the origin of rifling. Accordingly, the 
protection of the bore surface for even a short dis¬ 
tance in this region would be advantageous and it 
therefore seemed worth while to evaluate this device. 
Designs and performance are discussed in Chapter 32. 

15 5 Pre-Engraved Projectiles 

It soon became evident that the interaction be¬ 
tween the rotating band and the bore surface was an 
important source of deterioration of the surface. The 
stresses arising from the engraving of the band are 
likely to damage the surface, particularly by flatten¬ 
ing the lands; and the frictional heat superimposed 
upon that derived from the hot powder gases will 
obviously augment all adverse thermal effects. By 
using pre-engraved projectiles, that is ones on which 
“teeth” had previously been formed for engaging 
with the rifling grooves, these mechanical effects are 
obviated. 

Furthermore, a pre-engraved projectile was found 
to be a valuable research tool because it offers the 
opportunity of separating some of the effects due to 
the powder gas and those due to the rotating band. 
As the development proceeded, it became evident 
that the practical value of pre-engraved projectiles 
is much enhanced by using them in conjunction with 
a bore surface protected from the powder gases. The 
chromium plate is inert to the powder gases; and the 
pre-engraved projectiles do not cause mechanical de¬ 
terioration of the plate, as engraving-type projectiles 
do. The device, as finally put in the practical form 
described in Chapter 31, consisted of the combination 


CONFIDENTIAL 



STATUS OF PROBLEM IN 1946 


15 


of pre-engraved projectile and chromium-plated bore 
surface. 

1 - 5 - 6 Sabot-Projectiles 

A sabot offers a convenient means of using a light¬ 
weight projectile and therefore obtaining hyperveloc- 
ity in a standard gun. Sabot-projectiles do not repre¬ 
sent a new idea, having been tried out many years 
ago; but, because they offered a possibility of quick 
solution of the hypervelocity problem, it seemed worth 
while to make a serious attempt to develop a prac¬ 
tical projectile of this kind. Fundamentally, such a 
projectile consists of a subcaliber core, of either steel 
or tungsten carbide, supported in the bore by a sabot, 
which is a framework of some sort, including a base 
that fits the bore and carries a rotating band. Upon 
being fired, the projectile proceeds with the velocity 
appropriate to the weight of the assembly in relation 
to the powder charge. Shortly after leaving the gun, 
the base and auxiliary parts fly off by centrifugal ac¬ 
tion or air resistance, and the core goes on toward the 
target. It was recognized that a considerable amount 
of research and development would be necessary in 
order to meet various requirements such as stability 
of the projectile and in general the requisite accuracy 
of the shots. (See Chapter 29.) 

It is not immediately obvious that the sabot is a 
true solution of the hypervelocity problem, in the 
sense of providing higher velocities with reasonable 
gun life. Although it is a convenient method for com¬ 
bining large powder charges with projectiles of rela¬ 
tively light weight, it might appear at first glance 
that erosion would offer the same limitation as it does 
in conventional guns. It turns out, however, that the 
more noticeable variety of erosion in a gun, namely 
that near the breech end, depends more on the ratio 
of powder charge to bore surface than on the weight 
of the projectile. Consequently, with a lightweight 
projectile, the erosion may be expected to be much 
less than would be encountered with conventional 
guns and projectiles operating at corresponding muz¬ 
zle velocities. For this reason, such devices are prop¬ 
erly considered among the fundamental approaches 
to the practical realization of hypervelocity. 

15 - 7 Tapered-Bore Guns 

Another method for practical utilization of a light¬ 
weight projectile that has a sufficiently high sectional 
density is represented by the tapered-bore gun. This 


device was known to have beep developed in Germany 
prior to World War II. Remarkable results were 
credited to a 28/20-mm tapered-bore gun used by the 
Germans in the African campaign against the British. 
A projectile with flanges, or skirts, is reduced in diam¬ 
eter as it passes through the bore, of which all or a 
part is tapered so as to decrease in diameter in the 
direction of the muzzle. The further investigation of 
such guns clearly was a necessary part of the hyper¬ 
velocity program. Little was generally known about 
the principles of design of such projectiles for most 
effective performance, or of the performance of guns 
with the bore tapered in various ways. Results of the 
investigations by Division 1 on tapered-bore guns 
and projectiles will be found in Chapter 30. 

1,5,8 Other Devices and Processes 

There are several other approaches that were given 
consideration by Division 1 but were not actively ex¬ 
plored by that organization, although some of them 
are discussed in various chapters of this volume. A 
propellant of decreased erosiveness would obviously 
be of great advantage, for improved action might re¬ 
sult from lower flame temperature or lessened chem¬ 
ical activity. Furthermore, a powder of greater energy 
for a given bulk would simplify appreciably the de¬ 
sign of hypervelocity guns. Successful methods for 
cooling the bore surface or the entire barrel would 
counteract the tendency to erosion and would mini¬ 
mize other difficulties with high-power guns, espe¬ 
cially those operating at a high rate of fire. Inasmuch 
as a part of the troubles arising from high velocities 
are due to the action of conventional rotating bands 
on the bore surface, improved rifling and banding are 
desirable subjects for investigation. Still another in 
the list of devices for appraisal and investigation was 
the “booster” projectile, for which added velocity is 
given to the projectile near the end of its course by a 
rocket attached to the base. 

16 STATUS OF PROBLEM IN 1946 

1,6,1 Clarification of Advantages of 
Hypervelocity 

As one of the results of the comprehensive series of 
investigations undertaken by Division 1 and as the 
result also of the various tactical and strategic situa¬ 
tions occurring during the past 5 years, the advan¬ 
tages of hypervelocity now stand out in much sharper 


CONFIDENTIAL 



16 


THE PROBLEM OF HYPERVELOCITY 


relief than heretofore. (For an analysis of these ad¬ 
vantages, see Section 33.2.) In comparison with the 
attitude in 1941, there is now little doubt as to the 
reality of the advantages of projectiles with velocities 
in what we call the hypervelocity range, that is, those 
in excess qf 3,500 fps. 

For antitank warfare, for other ground use, for 
antiaircraft applications, or for aircraft armament, 
there is now little division of opinion as to need for 
improving conventional guns. Greater range, im¬ 
proved armor penetration, and increased probability 
of hitting a target, more especially a fast-moving air¬ 
plane, add up to an advantage that few will ignore. In 
1946, the problem is not so much to evaluate the 
merits of hypervelocities, but rather to find the best 
means of achieving them. h One would now indeed be 
conservative not to visualize the “gun of the future” 
as one firing projectiles with an initial velocity of 
more than 4,000 fps, and quite possibly more than 
5,000. It can be claimed fairly that the war period has 
clarified our notions on this subject. 

1,6,2 New Knowledge Concerning 
Nature of Gun Erosion 

The scientific groups who undertook to find quickly 
the cause and cure of gun erosion, who somewhat 
naively approached with confidence the task of mak¬ 
ing real progress in a field new and strange to them, a 
field in which significant advances had not been made 
for many years, have the satisfaction of knowing that 
they now have a picture of gun erosion that is clearly 
outlined, although not complete in all its details. 
They can visualize the several steps in gun erosion, 
and they can differentiate between the nature and 
probable cure of origin erosion on the one hand and 
muzzle erosion on the other. It is now quite definite 
that of the various possible factors in the erosion of 
guns, thermal effects, including melting, aided and 
abetted by chemical action, are the general regulators 
of origin erosion, and it is also apparent how other 
factors, such as stress and frictional heat associated 
with rotating bands, enter the erosion problem. In 
addition, a useful distinction has been drawn between 
the origin erosion of high-power, slow-fire guns where 
the overall effect is a loss of the material from the 
bore, and the deterioration of machine guns firing 
long bursts at moderate velocities where the outstand- 


h A statement of the British point of view on this subject 
was prepared in the summer of 1944. 395 


ing factor in bringing the gun to the end of useful life 
is a flattening of the lands on the softened bore surface. 

We now perceive that the erosion problem has sev¬ 
eral aspects, each of which in reality presents a sepa¬ 
rate problem. The first may be called the “high-power 
gun” or the “big-gun problem,” which is solved either 
by diminishing the heat to each unit of the bore surface 
or by supplying a liner or coating that is sufficiently 
chemically inert and at the same time is refractory 
enough to withstand, without melting, the intense 
heat applied over a short interval of time. The second, 
the “machine-gun problem,” finds its solution in the 
knowledge that here the desideratum ,is hot-hardness 
and in the use of a liner or a coating of adequate hot¬ 
hardness, or a combination of liners and coatings, in 
the protection of the bore surface. To these problems 
we may add a third, which relates to the cure of muz¬ 
zle erosion. Although in 1941 the extent to which 
muzzle erosion was prevalent in guns was not well 
known, and even its existence generally in guns was a 
subject for frequent debate, it is now known to vary 
in an orderly fashion with muzzle velocity and caliber. 
Indeed, its variation with muzzle velocity is well 
enough understood for us to predict confidently that 
at muzzle velocities of 4,000 fps or more it will begin 
to be of commensurate importance with origin ero¬ 
sion, and may even be the dominating factor. We also 
have sufficient understanding of the nature of muzzle 
erosion to afford promising leads for the avoidance of 
the difficulty. 

These various conclusions on the nature of erosion 
emanated partly from well selected tests and partly 
from basic research in a closely integrated series of 
investigations. 

1 - 6,3 Development of Several 

Solutions of the Hypervelocity Problem 

The investigations of Division 1 provided new 
knowledge that was not only interesting and sugges¬ 
tive but also of practical value, as is demonstrated by 
the fact that the results and conclusions led to the 
successful development of six useful, or potentially 
useful, solutions of the hypervelocity problem. At the 
time Division 1 discontinued its activities, various 
devices embodying these solutions either had been 
offered to the Services or were in an advanced state of 
development. These were the following devices al¬ 
ready described briefly in Section 1.5: The replace¬ 
able short steel liner, a variety of erosion-resistant 
liners and coatings, the pre-engraved projectile with 


CONFIDENTIAL 




STATUS OF PROBLEM IN 1946 


17 


chromium-plated bore, the Fisa protector with chro¬ 
mium-plated bore, sabot projectiles, and the tapered- 
bore gun with skirted projectiles. In the category of 
erosion-resistant liners are ones of hardened molyb¬ 
denum, which offers a gratifying solution of the high- 
power gun problem, and those of Stellite No. 21, 
which in the caliber .50 machine gun has given a solu¬ 
tion of the machine-gun problem that was successful 
to the point of being spectacular. Stellite-lined ma¬ 
chine guns were in large production at the close of the 
war. 

Other approaches to the practical attainment of 
hypervelocity were given some consideration by vari¬ 
ous civilian and military groups, but Division 1 did 
not participate actively in the investigation of those 
methods, including the ones mentioned in Section 
1.5.8. 


164 Demonstration of Successful Method 
for Attacking Difficult Ordnance Problem 

That the task undertaken by the Division 1 in 1941 
was one of formidable complexity, few will deny; and 
we think it fair to claim that the outcome was reason¬ 
ably satisfactory. It might even be claimed that 
whereas the group of investigators, upon first ac¬ 
quaintance with the field of gun erosion and hyper¬ 


velocity, found it a frozen subject, they left it a fluid 
one. The procedure that was followed might there¬ 
fore serve as a pattern for future investigations in the 
same fields or in ones of similar difficulties and scope. 
We feel that whatever success may have been attained 
was largely due to two factors. 

The first was the inclusion of an adequate amount 
of fundamental research. In the broad program of 
Division 1 neither research nor engineering develop¬ 
ment was neglected. There was at all times a judi¬ 
cious combination of effort in these two directions, 
the former being the fruitful source of new ideas and 
the latter the means to a useful end. For example, as 
soon as promising results were obtained with stellite 
liners, plans were put into effect for the production 
of caliber .50 machine-gun barrels with such liners; 
while at the same time further investigations were 
conducted in order to learn more about the funda¬ 
mental behavior of this material under various con¬ 
ditions. 

The second essential factor was the constant pre¬ 
servation of an open mind as to the possibilities of 
success of any development no matter what the ap¬ 
parent obstacles might be, and a refusal to be swayed 
unduly by unpromising results of various lines of 
attack in the past. The guiding philosophy might per¬ 
haps be described as a resolution to “lean neither to 
foolish optimism nor to unreasoning conservatism.” 


CONFIDENTIAL 







































































































































I 


PART II 
BALLISTICS 



Who should decide when doctors disagree 
And soundest casuists doubt like you and me. 

—Alexander Pope 
“Moral Essays” 


« 


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Chapter 2 

PROPERTIES OF POWDER GAS 

By F. C. Kracek a 


21 INTRODUCTION 

T he chemical system of powder and gas presents a 
complex in which chemical reactions occur at high 
speed with remarkable precision. The free energy of 
powder gas is much less than that of powder, and the 
reaction between them, when once initiated, goes for¬ 
ward rapidly—in common parlance, the powder burns. 
The gas is never in equilibrium with the powder; the 
powder must first be “vaporized,” the intermediate 
products then react further, and under suitable con¬ 
ditions, the final products come to equilibrium with 
each other as powder gas at a high temperature. 
These intermediate steps are of interest both theo¬ 
retically and in practice, for they lead to a better un¬ 
derstanding of the use of powder to give a push on a 
projectile or a rocket. 

Modern propellants are composed of certain prin¬ 
cipal constituents, explosive in nature, and other 
added constituents used as plasticizers, stabilizers, or 
solvents to modify the physical properties of the ex¬ 
plosives and make the resulting substance safe for 
handling as a propellant. The principal explosives 
used are cellulose nitrate (nitrocellulose, NC) and 
glycerol nitrate (nitroglycerine, NG); others more 
recently introduced are nitroguanidine (NQ) and hexa- 
methylene trinitramine (cyclonite, RDX). All are 
compounds of carbon, hydrogen, oxygen, and nitrogen, 
and contain enough oxygen in the reactive — 0 — N0 2 
(nitrate), or — N0 2 (nitro) groups for internal com¬ 
bustion. 

This combustion, the completeness of which de¬ 
pends upon the ratio of oxygen to carbon and hydro¬ 
gen, produces a quantity of heat, known as the heat 
of explosion, and results in the conversion of the solid 
powder to a gas composed principally of nitrogen, 
carbon dioxide, carbon monoxide, water, and hydro¬ 
gen. The relative proportions of these gases in the 
final products are determined by (1) the relative con¬ 
tents of carbon, hydrogen, oxygen, and nitrogen in 
the propellant, and (2) the temperature and pressure 
developed in the process of burning. Additional gas¬ 
eous constituents are present in the powder gas, as a 

a Physical Chemist. Geophysical Laboratory, Carnegie In¬ 
stitution of Washington. 


result of dissociative high temperature reactions, or 
of secondary reactions which occur on cooling. These 
are relatively unimportant, and will be considered 
only incidentally. 

It is important to note that explosives are inher¬ 
ently unstable at all temperatures, and can be pre¬ 
pared and kept only because the decomposition reac¬ 
tions proceed exceedingly slowly at ordinary temper¬ 
atures. The initiation of the decomposition requires a 
certain activation energy which may be supplied by 
heating or by impact. The resulting decomposition 
may be burning or detonation, the two being dis¬ 
tinguished by the speed with which the process takes 
place; the rate of detonation is about a thousand 
times that of burning. The detonation of explosives 
was studied intensively in other divisions of the 
NDRC; the burning, particularly of propellants, falls 
within the scope of the work of Division 1 and is 
treated in the following pages. 

Recent investigations have enabled us to put labels 
on the intermediate steps in the conversion of powder 
to the final products. The rates at which these steps 
occur are connected with the mechanism; some exper¬ 
imental evidence on this is also available. These prob¬ 
lems are treated in the second section of this chapter. 
The next section takes up the problem of equilib¬ 
rium, and evidence is presented to show that at least 
so far as the final reaction in powder gas is concerned, 
equilibrium is maintained among the final products. 

Internal ballistics is concerned with the conversion 
of the “latent” energy of the powder into mechanical 
energy through the medium of powder gas. The chem¬ 
ical thermodynamics of this process are presented 
briefly in Section 2.4. Experimental methods for de¬ 
termining the temperature of powder gas, discussed 
in the last two sections, present an experimental 
check on the theoretically calculated temperatures, 
and together with measurements of pressure, enable 
the ballistician to discuss the state of powder gas. 

22 THE BURNING OF POWDER 

221 The Overall Burning Process 

When a propellant is ignited in a closed chamber at 
a high density of loading, and suitable measurements 


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21 



22 


PROPERTIES OF POWDER GAS 


are taken, it is noted that soon after ignition is ini¬ 
tiated, the pressure begins to rise and then reaches a 
peak value in times of the order of milliseconds. Si¬ 
multaneously with the development of pressure, the 
contents of the chamber begin to emit radiation the 
intensity of which, like the pressure, reaches a peak 
value. 44 



MINUTES 

Figure 1 . Decomposition of nitrocellulose at pressures 
below 1.75 mm Hg. Reaction becomes self-sustaining at 
higher pressures. (This figure has been redrawn from 
Chart No. I in Ballistic Research Laboratory Report 
No. 353, Aberdeen Proving Ground.) 

The radiation at high densities of loading is contin¬ 
uous in the visible and near infrared regions of the 
spectrum; at very low densities of loading, discrete 
lines appear in the spectrum. These are due to in¬ 
organic constituents or impurities. The radiation at 
high densities is found to be not only continuous but 
also to have a distribution of intensities with respect 
to wavelength such as would be typical of a black 
body at a temperature T which is somewhat lower, 
because of the heat losses, than the adiabatic flame 
temperature To, characteristic of the propellant. As 


example, typical values of T 0 are 2650 K for NH, 
2440 K for FNH-M1, 3525 K for FNH-M2 powders. 
The temperature T has been identified as the tem¬ 
perature of the powder gas. 

The instants of time at which the peak of the radia¬ 
tion and the peak of pressure occur coincide closely 
and denote the conclusion of the process of burning. 
After the attainment of the peak values, heat contin¬ 
ues to be given off to the walls as the gas cools; in a 
gun, the gas is cooled not only by such heat transfer 
but also by conversion of a part of its thermal energy 
to kinetic energy of the projectile. The distribution of 
the energy of the powder is discussed in Section 6.5, 
in which the relation of the gas temperature to the 
energy is shown. 

The details of the burning of the powder cannot be 
analyzed in such experiments as described above and 
must be investigated in other ways, but the overall 
kinetics of the burning process can be learned from 
the measurement of the rate of pressure increase, at 
times preceding the attainment of maximum pressure. 

Experimentally, it is known that powders do not 
ignite until the surface is raised to a temperature near 
200 C; the self-sustaining reaction probably main¬ 
tains the surface at about 1000 K. 160 This burning- 
may be extinguished upon sudden reduction of pres¬ 
sure, as this causes appreciable cooling of the gas at 
the surface. This suggests that stable burning requires 
heat to be transferred to the surface by the hot gases 
in an amount sufficient to maintain the surface layer 
hot enough for rapid decomposition of the powder to 
take place. 

The amount of heat transferred to the surface may 
be expected to depend upon at least two factors; 
(1) the intensity and type of radiation which falls upon 
it, and (2) the number of molecules of high energies 
which collide with it per unit time. The part played 
by radiation, which is difficult to evaluate, will be 
discussed later. 

The number of colliding molecules will be nearly 
proportional to the pressure, but their average energies 
cannot be estimated without detailed knowledge of 
the reactions involved in the process of burning. If we 
assume, as a first approximation, that the entire reac¬ 
tion process takes place at the powder surface, then 
the gas molecules have energies corresponding to the 
flame temperature T 0 . On this assumption, the burn¬ 
ing rate may be expected to be high for powders 
whose flame temperature is high, and this is at least 
qualitatively confirmed by experiment. As we shall 
see, the reactions liberating the energy of the powder 


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THE BURNING OF POWDER 


23 


take place only partly at the surface, but mainly in 
the gas space, probably separated from the surface by 
a more or less thin layer of cooler gas. This reduces the 
rate of heat transfer but does not necessarily alter the 
relative order of burning rates for hot and cool pow¬ 
ders. 

2,2 2 The Stepwise Mechanism of Burning 

Before considering the rates of burning of powder it 
is desirable to present an outline of the reactions in¬ 
volved in the process of burning . b 

As long ago as 1907, it was demonstrated 457,458 
that an early step in the decomposition of nitrocellu¬ 
lose and nitroglycerine is the splitting off of N0 2 from 
the molecules of the explosives, and that the N0 2 is 
quickly reduced to NO when the gas is allowed to re¬ 
main in contact with the decomposing substance. Re¬ 
cent experiments 194 have shown that at very low 
pressures (below 1.75 mm Hg) the decomposition re¬ 
action of nitrocellulose is endothermic with the evolu¬ 
tion of gas, and may be stopped and restarted at will 
by controlling the heating (see Figure 1). At pressures 
above this value, the reaction becomes exothermic 
and hence self-sustaining; it proceeds with the evolu¬ 
tion of gas and sublimation of a white solid. At still 
higher pressures and with a chemically indifferent gas 
present (about 40 mm Hg), there is produced a resin¬ 
ous red liquid that gives the chemical reaction for 
aldehydes. This evidence indicates that the original 
cellulose nitrate molecule has been attacked, most 
probably at the points of attachment of the nitrate 
groups. 

The liquefaction of nitrocellulose has also been ob¬ 
served 350 at a pressure of about 50 mm Hg of nitro¬ 
gen, in experiments on controlled thermal decompo¬ 
sition. At 1-atmosphere pressure with air in the 
apparatus, liquefaction was not complete before an 
explosive inflammation occurred; the ignition was pre¬ 
ceded by an induction period which was shorter at 
higher temperatures. The activation energies deduced 
from the variation of the induction period with tem¬ 
perature, and of the rates of gas evolution during the 
quiescent preignition decomposition agree provision¬ 
ally with the activation energies for the splitting of 


b Most of the data upon which this outline is based has been 
accumulated during the war years by investigations at the 
Ballistic Research Laboratory, Aberdeen Proving Ground, 
194,198,243,548,549 contractors of Divisions 1 and 3, NDRC, 
107,159-163,547 an( j by investigators working for the British Minis¬ 
try of Supply. 334 ' 335 ' 345 * 350 ' 364 ' 396 


the — O —N0 2 bond, which is known from other data 
to be 40-50 kcal. 

The exothermic heat effect up to the point of con¬ 
sumption of the red substance has been measured to 
be near 500 cal/g, about one-half of the full heat of 
explosion of nitrocellulose. Similar experiments yield 
approximately the same values with double-base 
powders (rocket compositions, cordite S.C.). The gas 
produced at this stage contains the nitrogen of the 
powder principally as nitric oxide (NO); other gases 
present are C0 2 , CO, H 2 0, with very little H 2 and no 
N0 2 . Upon inflammation, H 2 definitely appears in the 



PSI 


Figure 2. Heat of explosion as a function of gas pres¬ 
sure showing the course of the curve in the region of the 
transition from nonluminous to luminous burning of 
propellant JP76. (This figure is based on Table VI of 
NDRC Report A-268.) 

gases, and the full heat of explosion tends to be re¬ 
alized. (See Figure 2.) 

Visual observation of the formation of flame from 
the burning of double-base rocket powder in nitrogen 
under pressure 161 shows a transition from nonlumi¬ 
nous to luminous burning at pressures near 300 psi. 
The flame, when first formed at the lowest pressure, 
is at some distance from the stick of powder, but ap¬ 
proaches it as the pressure is raised, as shown in Fig¬ 
ure 3. The transition from nonluminous to luminous 
burning in the gas phase marks the beginning of dis¬ 
appearance of NO from the gas. It is accompanied by 
an increase of the heat evolved with increasing pres¬ 
sure; the full heat of explosion is reached at about 700 
psi. At this pressure the NO in the gas virtually dis- 


CONFIDENTIAL 

































24 


PROPERTIES OF POWDER GAS 


appears. At the higher pressures the gas contains N 2 
and the water gas reaction gases CO, CO 2 , H 2 , and H 2 0 
in chemical equilibrium at the flame temperature. The 
gases of the nonluminous zone are not in equilibrium, 
and probably contain fragments of the original mole¬ 
cules ; thus C 2 H 2 has been detected in quenched gas 
from nonequilibrium burning. 107 The temperature of 
this zone has been estimated 0 to be about 1200 K. 



Figure 3. Nonluminous to luminous transition in the 
burning of double-base powder. (A) Luminous and non¬ 
luminous burning of powder HES 4016. (B) Variation 
in thickness of the dark zone, powder HES 4196 burn¬ 
ing in nitrogen (25 C) at pressures (left to right) 410, 
365, 310, 280, 255 and 225 psi. (These figures have ap¬ 
peared as Figures 1, 2, and 3 of NDRC Report A-268.) 

Summary 

The foregoing stages of burning may be summar¬ 
ized somewhat as follows: 

Stage A: Initial Decomposition. Through heat sup¬ 
plied from the outside or from subsequent reactions, 
the surface of the powder grain is brought to a tem¬ 
perature at which gases are just evolved at a rate de- 

c Personal communication from Capt. J. H. Frazer, BRL. 


termined by the heat supply. This endothermic stage 
may be the activation of decomposition of the — O — N 
bond. The gas may contain N0 2 . 

Stage B: Secondary Decomposition. The highly re¬ 
active gaseous products of Stage A react with one 
another or with the material on the surface of the 
powder, with the evolution of heat (which makes the 
reaction self-sustaining) and of further gases, and a 
volatile solid or aldehydic liquid, depending upon the 
pressure at which the reaction occurs. 

Stage C: Gasification of Intermediate Products. This 
reaction is a continuation of Stage B. It results in 
further evolution of heat and production of not fully 
reacted gases containing much of the nitrogen of the 
powder as NO, and the C, H, and O as C0 2 , CO, and 
H 2 0. Little H 2 is present. The net thermal effect of 
Stages A, B, and C is an evolution of about one-half 
of the full heat of explosion. It is probable that only 
part of the condensate of Stage B has reacted at this 
point. 

Stage D: Final Reaction. Oxidation of the remaining 
combustible products by nitric oxide, with the forma¬ 
tion of N 2 and of the water gas reaction gases, CO, 
C0 2 , H 2 , and H 2 0, in chemical equilibrium at the 
temperature of the flame, releases the second half of 
the energy of the overall reaction. It appears to be 
very sluggish, and is believed to have no effect on the 
rate of the overall reaction other than that due to the 
increased temperature of the powder surface. 

Stage E: Postburning Stage. After all of the energy 
of the powder has been liberated and the final prod¬ 
ucts evolved, the gas cools with a readjustment of 
the equilibrium of the water gas reaction, and pro¬ 
duction of secondary reactions, such as the formation 
of methane from CO and H 2 . These reactions occur 
with changes in energy which are small in comparison 
with the heat of explosion. 

Thus the steel of the bore surface of a gun is in con¬ 
tact with a variety of gases other than those found by 
an analysis of the relatively cool gases ejected at the 
muzzle. This situation, although vaguely suspected, 
had not yet been experimentally demonstrated when 
Division 1 began its investigation of gun erosion. It 
helps to explain the existence of some of the products 
of erosion found on the bore surface, which are de¬ 
scribed in Chapter 12. 

2 2 3 The Rate of Burning 

From the preceding discussion it appears that the 
rate of burning is controlled, not by the rate of forma- 


CONFIDENTIAL 








THE BURNING OF POWDER 


25 


tion of the final products in the gas phase (Stage D), 
but rather by the rates of the reactions taking place 
at or near the surface of the powder, Stages A, B, and 
C of the process of burning. Stage D contributes to 
the rate insofar as the heat liberated in the ^as phase 
reaches and penetrates through the surface and thus 
heats the solid powder. The thermal conductivity of 
the powder being low, and the times at disposal short, 
this penetration is exceedingly shallow. 

Rates of burning have been studied largely by 
measurements of the rate of change of pressure after 
ignition. Early experiments of this sort showed that 
the pressure increases slowly at first, both in closed 
chambers and in guns during the period before the 
projectile starts to move, and then more and more 
rapidly, the pressure-time curve somewhat resem¬ 
bling a hyperbola or a parabola. 

The powder is believed to burn at the surface in 
parallel layers, it being assumed that all the surface 
is ignited at the same time, and burns at the same 
rate. This rate may be defined as the rate r at which 
the parallel layers are removed in the direction nor¬ 
mal to the surface. The rate of “recession of the sur¬ 
face’ ’ may be related to the mass burned up to any 
instant by means of form functions derived from the 
geometry of the various types of granulations. 520 The 
functions proposed to express r in terms of the pres¬ 
sure P are of two general types, as given in equations 
(1) and (2). 

r s ^ = a + bP m (1) 

r = cPn (2) 

in which dx/dt is —W(df/dt ), W being the web thick¬ 
ness and / the fraction of the web which remains un¬ 
burnt at any time. The coefficient c in equation (2) is 
denoted B in the Division 1 system of ballistics de¬ 
scribed in Chapter 3. Equation (2), was proposed by 
Vieille in 1893. He recognized that the exponent n 
varies for different types of powder, having values 
from about 0.75 to 0.95. The rate equation (2) with a 
fractional exponent is difficult to handle in interior 
ballistics equations, and the general practice has been 
to adopt n = 1 and make a judicious adjustment in 
the value of c. 26 

For closed chamber firings, equation (1) has been 
rewritten 436 - 439 in the form of equation (1'). 

r = a + bP (1') 

This equation of a straight line with an intercept is a 
good approximation to the experimental results in 


the higher pressure range, abdjve perhaps 1,500 psi, as 
has been shown by measurements 141 for the range of 
pressure up to and beyond 20,000 psi. 

At lower pressures, such as are attained in rockets 
(600-6,000 psi) the exponential form, equation (2), is 
a better fit; in fact, with certain rocket powder com¬ 
positions, the exponent n is so small (about 0.5) as to 
rule out th£ linear form altogether. 161 At very much 
lower pressures, in the neighborhood of 1 atm, the ex¬ 
perimental curve shown in Figure 4 indicates a break 
in the curve of burning rate vs pressure, as if the 
mechanism of burning here suddenly changed. For¬ 
tunately this region is of no interest for ballistics. 



0 20 40 60 80 100 


PSI S 

Figure 4. The course of burning rate versus pressure 
for JP76 at different low pressures and temperatures 
showing the break at about 1 atm. (This figure has been 
based on Figure 19 of NDRC Report A-268.) 

In the intermediate pressure range (up to 2,000 psi) 
the rate of recession of burning powder was observed 161 
visually and photographically in a chamber equipped 
with a glass window. This direct measurement yields 
lower values for the burning rate than the indirect 
measurement from a pressure-time curve secured 159 
in a rocket motor. The discrepancy is ascribed to a 
more rapid consumption of powder in the rocket than 
in a closed chamber by removal of the products of 
combustion, and by erosion. 

A comparison of the burning rate rasa function of 
the pressure for the straight-line law with an inter¬ 
cept, and the simple power law is given in Figure 5. 
The data represented by the straight line are those 
for the closed chamber burning of a powder contain¬ 
ing 58% NC, 40% NG, 1% centralite, 0.2% diphen- 
ylamine, and 0.9% volatiles at a loading density of 
0.1. For the purpose of the comparison, the straight 


CONFIDENTIAL 


































26 


PROPERTIES OF POWDER GAS 


</) 

Ql 


3 

CD 



PRESSURE IN 1000 PSI 

Figure 5. Comparison of the course of the burning rate versus pressure for the linear law (with intercept) and the expo¬ 
nential law. (This figure is based on the data for Powder A-20 contained in Appendix II of OSRD Report 4382.) 


line is extrapolated beyond the 19,500 psi measured, 
to 50,000 psi. The curves representing the power-in¬ 
dex law are drawn torn = 0.7,0.8, and 0.9, respec¬ 
tively, adjusted to coincide with the straight line at 
20,000 psi. It will be noted that over the measured 
interval of about 1,000 to 20,000 psi, the error would 
not be great, whether the data were represented by 
the power-index law with n = 0.8 or 0.9, or by the 
straight-line law, although, over an extended range of 
pressure the four curves deviate from each other 
appreciably. 

A further comparison of the two burning-rate laws 
is afforded by the firings of the 3-in. Navy gun, 
and the 37-mm gun, T47 at the Taylor Model 
Basin. 39 ’ 65 131 132 (These guns are described in Sec¬ 
tions 4.2.2 and 4.2.3, respectively.) The burning rates, 


which have been discussed in two Division 1 re¬ 
ports, 113 * 116 are represented in Figure 6 as a function 
of pressure. It is seen that it would do considerable 
violence to the data to represent them by the pres¬ 
sure-index law. The values of the constants a (in. /sec) 
and b (in./sec/klb/in. 2 ) for equation (1') applied to 
these data are given in Table 1. The values of the 
constant a are somewhat erratic, and considerably 
higher than might be expected from the closed cham¬ 
ber firings on which Figure 5 was based. 

This anomaly is undoubtedly due to ignition effects. 
Localization of the early burning in a portion of the 
powder may result in a considerable local pressure 
increase and high burning rate at a time when the 
average pressure is still low. This effect is also evi¬ 
denced by a 40 per cent reduction in the value of a 


CONFIDENTIAL 















THE BURNING OF POWDER 


27 


Table 1. Values of burning-rate constants for powders 
used in 3-in. Navy gun and in 37-mm gun T47, as 
obtained from ballistic data. 113 


Powder* 

Class f 

in./secf 

in./sec/klb/in. 2 * 

[a-(256)/25]§ 

NH 


1.35 

0.260 

0.314 

Pyro 


1.97 

.242 

.321 

Ml 

1 

2.59 

.239 

.343 

Ml 

4 

1.57 

.255 

.318 

M5 

6 

2.01 

.458 

.538 

M5 

10 

1.49 

.483 

.543 

M5 

12 

1.85 

.451 

.525 


* NH and Pyro were fired in the 3 in. 
t Classes: (1) short primer, 1.62-lb projectile (medium) 

(4) long primer, 1.62-lb projectile (short) 

(6) short primer, 1.62-lb projectile (medium) 

(10) short primer, 1.34-lb projectile (short) 

(12) short primer, 1.92-lb projectile (medium), 
t From Reference 113. 

§ The last column gives the “effective” 6 at 25 klb for comparison with 
Hirschfelder’s c referred to in the text. 


in the firings of the 37-mm gun, resulting from the 
use of a long primer that favored even ignition. The 
constant b has about the same value for all three 
single-base powders (0.250 + 0.007), and is nearly 
80 per cent greater for the double-base M5 powder. 

The use 26 of the approximation of c = 0.36 for 
single-base and 0.51 for double-base powder in the 
first power burning law without intercept (Section 
3.2.5) can be reconciled 113 with these data when it is 
noted that the rather large term in a compensates for 
the lower values of b ; the rates given by the two laws 
agree to within 10 per cent at 25 klb/in. 2 . In actual 
use, the constants in the Division 1 system of ballis¬ 
tics would be adjusted by fitting the firing data for 
either the maximum pressure, or muzzle velocity. 


2 2 4 Effect of Various Factors on the 
Rate of Burning 

Temperature of the Powder 


Powders burn at a rate which causes the surface to 
recede so fast that very little heat is conducted below 
the surface and the bulk of the powder, therefore, 
does not change in temperature during the burning. 
The initial temperature of the powder has been found 
to increase the rate by 3 to 6 per cent for every 10 
centigrade degrees rise in the powder temperature. 
For burning in rockets, the effect of the initial tem¬ 
perature could be expressed 163 by equations (3 and 4) 


a'-+ b'P 

t\ tp 
c ' p n 
t\ tp 


(3) 

(4) 


where t p is the (centigrade) temperature of the pow¬ 
der; ti is a constant whose value is typical of each 
powder, and varies from 200 to over 300 degrees for 
rocket compositions, being larger for powders show¬ 
ing smaller dependence of the rate on temperature. 



/ 

/ 

/ 

/ 

y 

y- 

y n c 

S 

D 


/ 

/ 

/ 

/ 

/ 

/ 

p '* • • 



\ 

* o \ 


" • • 

• 



* 

y/n 

X 

3" GUN PYRO POWDER 
o 100% CHARGE 
•75% CHARGE 
x 50% CHARGE 


V F0 

I_ 

R COMPARISON 
(AVERAGE) 

- Ml POWDER 37MM T47 GUN 

- M5 POWDER 37MMT47 GUN 


10 20 30 

PRESSURE IN 1000 PSI 


Figure 6. Burning rates in 3-in. Navy gun and 37-mm 
gun T47, for different powders showing the inapplicabil¬ 
ity of the exponential rate law to gun results. 


Composition of the Powder 

The burning rates are higher for powders which lib¬ 
erate larger amounts of energy on burning. The 
meager data available indicate that for burning at 
gun pressures, the burning rate constant b is increased 
0.03 to 0.05 (in./sec/klb/in. 2 ) for every 100 Kelvin de¬ 
grees rise in the adiabatic flame temperature. More 
consistent burning rate data are needed, both for 
closed chambers and guns. 

Volatile constituents of powders serve as coolants, 
and lower the burning rate by some 10 per cent for 
every 1 per cent of solvent and moisture present. 162 


Radiation 

An increase in the radiation received by the powder 
increases the rate of burning. This has been shown 
for firings in a pressure vessel with reflecting walls. 
Previous experiments 160 had shown that a double¬ 
base rocket powder (40% NG) burning in open air 
emits strong radiation, possibly due to chemilumines¬ 
cence from vibrationally excited molecules of C0 2 
and H 2 0, at 2.8 and 4.4 ju, superimposed on weak 
black body radiation. In a pressure vessel with a 


CONFIDENTIAL 























28 


PROPERTIES OF POWDER GAS 


quartz window, the 2.8 ju band was found also at 2,000 
psi pressure (the 4.4 /x band was blanked out by the 
opacity of the window). 

Potassium salts incorporated in powders have been 
known to increase the burning rate. When powders 
containing several per cent of potassium salts were 
burned in air at atmospheric pressure, a strong emis¬ 
sion was found at the red line of potassium, 7660 A. 
The black body radiation intensity was negligible in 
these experiments. Burning the powders in the pres¬ 
sure vessel with a window, with freshly polished walls, 
at pressures up to 1,600 psi, the burning rate was in¬ 
creased about 30 per cent over that observed with 
absorbing surroundings. Sodium salts did not seem 
to have such an effect. 

It is not known what effect this discrete radiation 
has under gun conditions, but it must be greatly sub¬ 
ordinated to that of the black body radiation, inas¬ 
much as all the discrete lines in the visible and near 
infrared spectrum merge into a continuum of emis- 
sivity approaching unity at density times path- 
length of the order of 0.06 g per sq cm. 44 

The radiation, regardless of its nature, if it pene¬ 
trates into the powder, gradually raises the powder 
temperature and the burning rate. This effect is es¬ 
pecially important with transparent powder near the 
end of burning when the interior of the web may be 
heated by radiation from both sides. In rockets, this 
may cause a marked progressiveness of the rate of 
burning and a sharp upward rise in pressure near the 
tail end of burning. It may be counteracted by incor¬ 
porating darkening agents in the powder, such as 
lampblack or nigrosene dye which prevents the radia¬ 
tion from penetrating to any depth into the powder. 

Little attention has been given to radiation effects 
in guns. It may be mentioned that irregularities of 
ignition have been counteracted by treating the pow¬ 
der with an aqueous solution of paraphenylenedia- 
mine which blackens the surface uniformly; if an al¬ 
coholic solution is used, the blackening penetrates 
throughout the body of the powder, and the burning 
rate may rise dangerously high. 243 These effects un¬ 
doubtedly arise at least partly from absorption of 
radiation at the surface. 

Further investigation of this phenomenon is highly 
desirable, because of its possible application to hyper¬ 
velocity guns. d In the past the density of loading for 
cannon has been limited to about 0.7 because the use 

d Such an investigation had been planned by Division 1, 
NDRC, in connection with its development of a hypervelocity 
90-mm gun, described in Chapter 33. 


of higher densities has resulted in irregularities in ig¬ 
nition. The use of blackened powder combined with a 
long primer may represent a simple means of over¬ 
coming this difficulty and permitting the use of the 
higher densities of loading that are essential if hyper¬ 
velocities are to be achieved with existing propellants. 

2 2-5 Stepwise Mechanism in Relation to 
Rate of Burning 

According to the stepwise mechanism of burning 
the gases resulting from the incomplete burning in 
Stages A, B, and C accumulate in the space around 
the powder and are ignited in Stage D with heating of 
the products to the flame temperature. Although the 
average rate of burning of the powder shows no sud¬ 
den change in the corresponding region of pressure, 
there is evidence that local irregularities in the rate 
occur. 

High frequency variations in dP/dt have been ob¬ 
served, especially with perforated powder, beginning 
at pressures of the order of those at which the appear¬ 
ance of visible flame has been observed 161 or at 
slightly higher pressures. This so-called “hash” on 
the oscillograph trace may be due to earlier ignition 
in the perforations caused by local pressure excess. 
Experiments 198 in which dP/dt was directly measured 
as a function of time with unperforated, very smoothly 
burning powder showed that the dP/dt vs t curve ex¬ 
hibits a hump in the corresponding pressure region, 
as shown in Figure 7. The irregularity is too small to 
be observed in the curve of P vs t. The irregularity in 
dP/dt is believed to be the result of the sudden igni¬ 
tion of the intermediates to form the final products at 
the beginning of Stage D. 

23 CHEMICAL EQUILIBRIUM 

IN POWDER GAS 

2,3,1 Reactions During the Cooling 
of Powder Gas 

Questions arise as to the state of the powder gas 
both while burning is in progress and afterwards. If 
equilibrium among the various constituents prevails, 
the composition may be expected to vary with the 
temperature according to the thermodynamic laws. 
Because of the heats of reaction, this modifies the 
energy relationships; other properties of the gas also 
vary with the compositions. All these questions have 
a bearing on the calculations of interior ballistics. As 


CONFIDENTIAL 




CHEMICAL EQUILIBRIUM IN POWDER GAS 


29 



-30 -20 -10 0 


MSEC 

Figure 7. Comparison of experimental (circles) and average calculated dP/dt versus time curves for a very smoothly 
burning cord powder. Arrow indicates the conclusion of the irregularity in the burning rate. (This figure appeared as 
Figure 20 in Ballistic Research Laboratory Report No. 456, Aberdeen Proving Ground.) 


will be shown in this section, there is reason to believe 
that equilibrium is maintained in the gas, at least 
during the postburning period. Such a conclusion is, 
of course, anticipated by the success with which in¬ 
terior ballistic calculations predict the behavior of 
powder in guns. 

The principal chemical reaction among the gases 
in a gun after the powder is burnt is the readjustment 
of the concentrations of the water gas constituents. 
During the burning the energy liberated heats the 
products to a temperature somewhat below the adi¬ 
abatic flame temperature, the deficiency being ac¬ 
counted for by heat losses. As the gas cools after¬ 
wards, the water gas reaction components continue 
to react with each other and if the rate of this reaction 
is high enough to maintain equilibrium, the relative 
composition at any temperature T is given by the 
equilibrium “constant ’ 1 

KiGi = GmC ™ = K (A,T) (5) 

^co 2 u h 2 

where Ki is the equilibrium constant for the reaction 


in terms of ideal gases and the factor G\, which is 
unity for ideal gases accounts for the gas imperfec¬ 
tion at the prevailing density. 24 The equilibrium con¬ 
stant Ki increases with the temperature, and its 
values have been tabulated. 

When powder gas after cooling to room tempera¬ 
ture is analyzed, the relative concentrations of the 
water gas constituents do not correspond to a value 
of K\ characteristic of room temperature, but rather 
to one for a higher temperature, which is of the order of 
1000 K or higher;the observed constant KiG\ depends 
upon the rate at which the gas was cooled and de¬ 
creases with increasing density of loading, as may be 
shown from data 443 secured in massive chambers. 
This indicates that at the temperature corresponding 
to the observed value of KiGi the rate of reaction had 
become negligible in comparison with the rate of 
cooling, and the observed gas composition had been 
frozen-in. (See also Section 2.3.4.) 

Numerous analyses of powder gases have been per¬ 
formed, with little attention paid to the rate of cool¬ 
ing; the values of the equilibrium constant calculated 


CONFIDENTIAL 














30 


PROPERTIES OF POWDER GAS 


from such analyses are of little significance for our 
purpose. The first systematic study 451 of the influence 
of rapid cooling on the gas composition showed that 
powder gas cooled in an explosion vessel chilled to 
—120 C contained more CO than gas in a vessel 
heated to 90 C, and also that powder gas quickly ex¬ 
panded from a rifle whose barrel was cut off at various 
lengths contained more CO near the origin of rifling 
than at the muzzle 452 where the gas arrived cooler. 
These changes of composition will be seen to corre¬ 
spond to the expected changes in the freezing-in of 
equilibrium under the stated conditions. 


2 3 2 Quenching of Powder Gas Reactions 

An extended series of experiments on the freezing- 
in of equilibrium in powder gases, or “quenching,” as 
it is more conveniently called, was performed for 
Division 1 at the Geophysical Laboratory. 107 In these 
experiments an effort was made to obtain as rapid 
cooling as possible, by expanding the gases from a 
small firing chamber through a nozzle into a much 
larger expansion chamber. The expansion ratio was 
about 230:1. 

If the nozzle performed adiabatically, the gas 
would first be cooled on entrance into the throat in 
the ratio T t /T c = 2/(y — 1) where T t and T c are the 
temperatures of the gas in the throat and the firing 
chamber, respectively, and y is the ratio of the 
specific heats of the gas. On the exit side further cool¬ 
ing by expansion into an evacuated chamber should 
reduce the temperature momentarily in the ratio 
T/T t = ( P/P t ) 7/(7-1) where the subscript t refers to 
the values in the throat of the nozzle. Unfortunately, 
this momentary cooling is not permanent, as the gas 
reheats in the expansion chamber by recompression 
to the final low pressure (about 1 atm). In the mean¬ 
time, however, much heat is given up to the walls, 
and the net effect is a relatively rapid cooling. Tt will 
be instructive to note that the walls of the chamber 
that was used should absorb about 0.2 to 1 cal/cm 2 
(depending on the energy of the powder) per gram of 
powder fired, in the time the firing chamber is emptied, 
in order to cool the gas quickly to about 1500 K. 

In a second method of quick cooling, the incoming 
powder gas was mixed with an inert gas (usually ar¬ 
gon) in the expansion vessel. This permitted the heat 
to be distributed and delivered to the walls more 
slowly. The momentary temperature attained by the 
mixture before any heat is given up to the walls is ex¬ 


pressed by 

T _ niCi Ti + ?teC2 T 2 , 

riiCi + rwh 

where n, c , and T are the number of moles, the heat 
capacity, and the temperature of the gases which are 
distinguished by the subscripts 1 and 2. The method 
therefore is an attractive one; however, because of 
imperfect mixing, its full value is not realized in prac¬ 
tice. 

The pressure at which the powder gas was released 
for expansion was controlled by using rupture disks of 
selected thickness to close off the exit of the firing 
vessel; it varied from 500 to 1,850 atm. 

A large variety of powder was used, ranging from 
single-base NH and FNH and cool RDX powders at 
one extreme to very hot, double-base powders such as 
Hercules ballistite in fine granulation at the other, T 0 
ranging from 2420 to 3860 K. With powders of low 
flame temperature, the quenched gas had composi¬ 
tions yielding values of K\ of the order of 3.5 to 5 e cor¬ 
responding to quenching temperatures of 1700 to 
2000K for experiments in which nearly all the powder 
was burnt before breaking of the rupture disk. Gas 
from powders of very high flame temperature usually 
gave lower values of K h about two to three, which 
may be interpreted to mean that the additional heat 
delivered into the expansion chamber by the gas 
could not be dissipated as rapidly as in the previous 
case, so that the gas cooled more slowly and had its 
equilibrium quenched at a lower temperature. 

For both these extremes, there was little difference 
in the results whether the gas was expanded into an 
evacuated chamber, or into a chamber filled with ar¬ 
gon. On the other hand, with fine silver coils in the ex¬ 
pansion chamber as cooling agent, quenching occurred 
more efficiently in expansions into a vacuum than 
into an argon atmosphere. For example, with an NH 
powder, Ki (vacuum) was 4.9, where K\ (argon) was 
4.3, corresponding to temperatures of 2040 and 1870K, 
respectively. The lower value observed in argon 
may have been due to imperfect mixing of the two 
gases. 

When the amount of powder fired exceeded appre¬ 
ciably that to produce rupture, the results differed 
according to conditions. As outlined in the next para¬ 
graph for the NH powder, T 0 =2651 K. 

e The equilibrium was quenched always on the exit side of 
the nozzle where the pressure was of the order of 1 atm and 
where therefore the gas imperfection was negligible, so that 
A'iG'i = K\, the value determined for ideal gases. 


CONFIDENTIAL 





CHEMICAL EQUILIBRIUM IN POWDER GAS 


31 


Pressure at Rupture: Above 1,000 Atm, 

NH Powder 

At these pressures, the observed values of Ki did 
not differ appreciably, whether the amount of powder 
loaded and burnt just sufficed to produce rupture, or 
whether an excess of powder was used; in either case, 
they averaged about 3.8. However, with excess pow¬ 
der, somewhat less of the carbon of the powder than 
before was accounted for in the gas as CO and C0 2 , in¬ 
dicating a tendency toward less complete burning. 
Often acetylene, and in some cases, small quantities 
of nitric oxide were found in the quenched gas. 

Pressure at Rupture: 500 Atm, NH Powder 

Three different regions of loading density could be 
distinguished when the pressure at rupture was 500 
atm. 

Region 1 . Powder all burnt at rupture (0.6 g, caliber 
.30 nozzle, A 0 = 0.061). Values of Ki were near 5.0, cor¬ 
responding to a quenching temperature somewhat in 
excess of 2000 K. This is believed to be a normal value, 
as the quantity of gas is small, and heat should there¬ 
fore be dissipated quickly, the process giving rise to 
efficient quenching. 

Region 2. Powder not all burnt at rupture (0.6 to 
> 3.0 g powder, caliber .30 nozzle, A 0 = 0.061 to 0.37). 

1. With argon in the expansion chamber the values 
of K\ obtained were anomalously high (about 8 or 
more) and corresponded to quenching temperatures 
in excess of the flame temperature. (See Figure 8.) 
Although little or no unburnt powder was recovered 
from the apparatus, the carbon of the powder ac¬ 
counted for in the gas was uniformly low, often below 
70 per cent. The percentage of H 2 in the gas was very 
low. C 2 H 2 and much NO were present in the gas, NO 
often accounting for half of the nitrogen. Tarry resi¬ 
due accumulated in the apparatus. The powder con¬ 
sumed after rupture evidently did not burn com¬ 
pletely by the reactions of the last stage of burning, 
but rather decomposed only partially beyond the 
stage of the red liquid and nitric oxide. This interpre¬ 
tation is favored also by the fact that the small per¬ 
centage of H 2 , commonly found in incomplete burn¬ 
ing, was largely responsible for the high values of Ki 
determined, for theCO/C0 2 ratio usually had a slightly 
lower value than in normal, complete burning. 

2. When the expansion chamber was originally 
evacuated and the pressure at release was 500 atm, the 
values of K\ were somewhat high, but not as anoma¬ 


lous as when argon was preserit in the chamber; how¬ 
ever, C 2 H 2 and NO were present, and the excess pow¬ 
der burned incompletely. It is believed that the more 
anomalous burning found in the presence of argon is 
the result of local quenching in the gas mixture, argon 
acting as diluent. 

Region 3. Powder not all burnt at rupture, (3.5 g 
powder or more, caliber .30 nozzle, A 0 greater than 
0.37). In expansions either into argon or a vacuum, 
these charges gave normal values of K h and both 


12 _ QQ --Q. 


u 

X. £ 

</> 

o 

" D 

o 

2*1 O 

2 — 

13.6 




C 

> O < 

> | 

1 

1 



J 

O 


1 

1 

1 

1 

1 

1 

1 


O 

1 

l 

' 

1 

1 

1 

1 

1 

1 

O ARG 
• VACl 

ON 

JUM i 

| 

1 

1 


GRAMS POWDER FIRED PER 9.5 CU CM CHAMBER 


Figure 8. Values of the water gas reaction constant 
K\Gi for NH powder gas quenched by pressure release 
at 500 atm into argon (open circles) and into a vacuum 
(full circles), with increasing excess charge of powder. 


C 2 H 2 and NO were absent from the gas. The C, H, 
and O of the powder were fully accounted for in the 
gas, so that all indications point to a complete burn¬ 
ing of the water gas constituents and N 2 . The appar¬ 
ent explanation for this sharp return to normal burn¬ 
ing with release at 500 atm pressure and a very large 
excess of powder loaded is that, at these increased 
densities of loading, the friction encountered in the 
nozzle by the increased volume of the products main¬ 
tains a back pressure in the ignition chamber high 


CONFIDENTIAL 















32 


PROPERTIES OF POWDER GAS 


enough for the rate of burning to remain above the 
value for incomplete burning. 

Following the instant of rupture, the pressure un¬ 
doubtedly falls sharply from the 500 atm value at 
rupture to a lower value; in view of the observations 
on the variation of the rate of burning discussed in 
Section 2.2.5, the minimum pressure reached must not 
be much lower than about 60 atm if the final stage of 
burning is to be maintained. This means that soon 
after the initial drop in pressure a steady state must 
be established in the nozzle such that the rate at 
which discharge takes place is balanced by the rate at 
which gases are produced. 163 On this explanation, the 
steady state is never quite reached in Region 2 of 
loading densities. 

The anomalous values of K\ at releasing pressures 
of 500 atm were not obtained with fast-burning pow¬ 
ders of moderate and high flame temperatures. It may 
be suggested that, at rather low releasing pressures, a 
Region 2 of loading densities would also be realized 
with such powders. At such low pressures, the burning 
rate would be low at release and subsequent readjust¬ 
ment of the rates of burning and of discharge would 
allow the pressure to fall below the critical value for 
complete burning. This contention is supported by ex¬ 
periments performed without using rupture disks, the 
burning powder evolving gases freely into the expan¬ 
sion chamber without obstruction. Argon at 1 atm 
pressure was used to provide the initial pressure nec¬ 
essary for ignition. In such experiments, all powders 
gave values of K\ in excess of 10, sometimes in excess 
of 50; H 2 concentration in the gas was always low, that 
of NO was high, and consumption of the carbon of the 
powder only near 50 per cent. 

2 3 3 Occurrence of Methane, Acetylene, 
and Nitric Oxide in Quenched Gas 

Methane is usually assumed to be formed in powder 
gas as the result of secondary reactions, such as for 
example, CO + 3H 2 = CH 4 Hr H 2 0, or 2CO + 2H 2 
= CH 4 + C0 2 . Acetylene has not been systematically 
reported by other investigators. In the quenching ex¬ 
periments described in Section 2.3.2, chemical tests 
showed that when nitric oxide was present in the gas, 
the hydrocarbon constituents in the quenched gas 
contained acetylene, but when nitric oxide was absent, 
acetylene also was usually absent. 

The nitric oxide was present, as already mentioned, 
in experiments with excess powder, and particularly 
when the pressure at release was low. It is believed to 


be the result of primary decomposition in the early 
stages of burning. The presence of acetylene in the 
same experiments indicates that it likewise is a prod¬ 
uct of the early stages. Neither acetylene nor nitric 
oxide formed under such conditions would be expect¬ 
ed to survive the high temperatures prevalent in the 
final stage, for thermodynamic considerations show 
that both these gases, if formed at high temperatures, 
would tend to decompose in a cooling gas. Methane 
on the other hand is expected to form in increasing 
quantities as the temperature falls. This, of course, 
does not preclude methane being formed also as a 
product of the primary decomposition. Some evidence 
on this point is given in Section 2.3.5. 

2 3 4 Water Gas Reaction 

Equilibrium in Quenching 

The foregoing discussion of quenching experiments 
gives no indication whether the concentrations of the 
water gas reaction components determined corres¬ 
ponded to equilibrium values for the reaction at the 
quenching temperatures. To obtain this information, 
it is necessary to compare the experimental results 
with values expected from the powder composition. 107 
Such a comparison can conveniently be made in the 
following manner. 

The expected gas composition can be calculated 24 
from the powder composition converted to moles of 
atoms of C, H, O, and N per unit weight of powder, 
and an assumed value of KiGi (= Ki at ordinary pres¬ 
sure) corresponding to some chosen gas temperature. 
The values of CO/C0 2 were evaluated for Ki = 3.5 
(T = 1687 K) for the various powders studied, and 
plotted against the adiabatic flame temperature, as 
in Figure 9. In order to make the desired comparisons 
with the quenching data, the observed gas composi¬ 
tions from the quenching experiments likewise were 
recalculated to K\ = 3.5 from whatever value of K\ 
given by the experiment, and the corresponding 
values of the CO/C0 2 ratios were evaluated. Only 
data in which the evidence pointed to complete burn¬ 
ing of the powder were utilized for these comparisons. 
Thus, none of the experiments for which K i was 
anomalously high, or in which appreciable concentra¬ 
tion of NO was found, were included. 

The plot of observed and expected values of the 
CO/C0 2 ratios vs the flame temperatures is given in 
Figure 10. It is seen from this figure that the agree¬ 
ment between observed and expected is fair, especially 
when it is remembered that (1) the nominal powder 


CONFIDENTIAL 



CHEMICAL EQUILIBRIUM IN POWDER GAS 


33 


compositions do not always represent the actual 
powder compositions, by reason of variations in the 
volatiles, etc., (2) the efficiency of quenching is only 
moderate, and (3) the gas analysis, in the hands of 
even the best experimenters 443 makes an accounting 
which may vary as much as 10 per cent from the 
mean for a series of experiments on the same powder. 
It may, therefore, be concluded that within the ex¬ 
perimental error, the concentrations of the water gas 
components readjust themselves with the tempera¬ 
ture as rapidly as the temperature changes occur 
down to the instant of quenching. 

2 3 5 Equilibrium Among the Carbon Atoms 

As we have already seen, powder gas may contain 
products of incomplete combustion or of secondary 
reactions in addition to the water gas reaction con¬ 
stituents. With the discovery of induced radioactiv¬ 
ity, a direct means has become available for testing 
the distribution of any one element among the dif¬ 
ferent gaseous species of which the element is a con¬ 
stituent, by using its radioactive isotope as a tracer. 
From such experiments conclusions may be drawn 
concerning the state of equilibrium in the gas as a 
whole. This can be done for carbon by using prefer¬ 
ably its radioactive isotope (C 14 ) which,is long lived 
and hence lends itself particularly well for use as a 
tracer element. 

The principle of such an experiment is portrayed 
by Figure 11. The powder charge is coated with a 
small quantity of a dissociable carbon compound con¬ 
taining the radiocarbon, as indicated by the red lines 
outlining the powder grains in the upper part of the 
figure. The carbon atoms in the powder itself are not 
radioactive. When the powder is burned, the radio¬ 
carbon atoms in the coating are converted to gas 
along with those from the powder grains. Exchange 
reactions take place among them and some of the 
resulting carbon monoxide and carbon dioxide mole¬ 
cules become radioactive from the presence of tracer 
atoms of radiocarbon as shown by the red circles in 
the lower part of the figure. The proportion of carbon 
monoxide and carbon dioxide molecules that become 
radioactive can be determined by recovering the 
powder gases and measuring their specific activity. 

The first experiment 244 of this sort, performed at 
Aberdeen Proving Ground, demonstrated that tracer 
atoms (of heavy carbon C 13 instead of radiocarbon 
C 14 ) were uniformly distributed between the mole¬ 
cules of carbon monoxide and carbon dioxide re¬ 


covered from a caliber .30 rifle after firing. This 
result indicated that these two gases had been in 
mutual equilibrium in the powder gas. 

Later a more extensive investigation 61 was under¬ 
taken with the quenching apparatus referred to in 
Section 2.3.2, with radiocarbon (C 14 ) as tracer added 
to the powder. When the conditions of burning were 
such that the powder burned to the water gas con¬ 
stituents and nitrogen, the earlier results with respect 
to carbon monoxide and carbon dioxide were con¬ 
firmed. It was found in addition that methane present 
in powder gas resulting from complete burning is in 
equilibrium with the CO and C0 2 , and therefore, 
must be the result of secondary reactions after burn¬ 
ing. 

In experiments with incomplete burning at the 
instant of release of pressure (compare Section 2.3.2), 
none of the carbon gases were in mutual equilibrium. 
The methane present was essentially inactive, and 
therefore, presumably a product of primary decom¬ 
position of the organic residues. The acetylene al¬ 
ways carried about a half of the proportionate activ¬ 
ity of CO or C0 2 , which would indicate that it was a 
product of CO or C0 2 with a hydrocarbon fragment 


6079£T 
6084 \ 


A RDX 

C) 

IG) 


\ 

V 6080 

\ 

A IMR 

A Ml (N 
JO NO 

O M2 (N 


6081 




\ 

6040c^ 



17 a 

6065A\ 



\ 

\ 

\ 




\ A 15 

\A36I3 6082^ 



28' 

\ 

99 A ^4690 

\ ^ 6083 



no/t^ aX 

22825 > 

15 

V 




\ 

\ 

\ 

280540\ 

\ 




240 0 

\ 

1 

019 




\ 

\ 

\ 

\ 

_2£l 


2000 2500 3000 3500 4000 


Tn. IN DEGREES K 

Figure 9. Values of CO/CO 2 in the gas from various 
powders at K\G\ =3.5, calculated from powder analysis, 
and plotted versus the adiabatic flame temperature. 


CONFIDENTIAL 


























34 


PROPERTIES OF POWDER GAS 



T a IN DEGREES K 

Figure 10. Values of CO/CO 2 in the gas from various powders at KiGi=3.5, versus the adiabatic flame temperature, 
comparing values calculated from powder analysis with values deduced from the quenched gas compositions. The calcu¬ 
lated values are the same as those shown in Figure 9. 


CONFIDENTIAL 

































THERMODYNAMICS OF POWDER GAS 


35 


formed in the primary decomposition of excess pow¬ 
der already stripped of its radiocarbon. The lack of 
mutual equilibrium of CO and C0 2 in the incomplete 
burning is the result, most probably, of the excess pow¬ 
der being decomposed at low pressure to the products 
of the early stages of burning described in Section 
2 . 2 . 2 . 


thermal and chemical equilibrium. The agreement be¬ 
tween the calculated and observed temperatures, 
therefore, indicates that the full calculated energy 
of the powder is liberated by an equilibrium process 
in the high-pressure burning. 

24 THERMODYNAMICS OF POWDER GAS 


2 3 6 Equilibrium in Powder Gas— 
Conclusions 

The definite conclusions concerning the state of 
equilibrium in powder gas that can be given at this 
time are the following. 

The agreement of the observed and expected values 
of C0/C0 2 ratios from quenching experiments with 
complete high-pressure burning indicates that changes 
in chemical equilibrium among the water gas con¬ 
stituents closely follow the changes of temperature 
during the postburning stage, down to about 1500 K 
at pressures above 60 atm. 

The carbon gases are in mutual equilibrium in high- 
pressure burning, both during the active-burning and 
the postburning stages. 

When burning occurs at reduced pressure, equilib¬ 
rium is maintained only when a certain minimum 
pressure (about 60 atm) is exceeded; below this, 
water gas reaction equilibrium is not established. 
This is shown by the results of quenching experi¬ 
ments both with and without tracer radiocarbon in 
the powder. 

One more piece of evidence may be noted here. 
Measurements of the temperature of the powder gas 
show 44 that when a reasonable allowance is made for 
heat losses, the measured temperature closely agrees 
with the calculated adiabatic flame temperature. This 
temperature is calculated by thermodynamic meth¬ 
ods from known energy data, and for products in 


2,4,1 The Thermodynamic Problem 

When powder burns, a certain amount of energy is 
liberated by the combustion reaction. In a closed 
vessel, if the process occured adiabatically, this 
energy would heat the gas to the adiabatic flame tem¬ 
perature T 0 . In the actually realizable process, some 
of the energy is expended in heating the chamber, 
and, in the case of a gun, in doing external work. As a 
result, the energy available for heating the gas is less, 
and the final gas temperature T is lower than the 
adiabatic temperature TV The energy lost by the gas 
to the surroundings may be termed the energy re¬ 
leased. Thermochemically, it is the energy stored in 
the powder (per gram) less the energy stored in the 
gas in its final state of temperature T, density A. 

The thermodynamic problem is, then, to evaluate 
these energies and the related thermodynamic func¬ 
tions such as the heat content, the entropy, and the 
heat capacities for the actual gases and their mixtures 
as they occur in powder gas. Since the properties of 
the mixture depend upon the composition, and this in 
turn depends upon the equilibrium of each component 
reaction, the equilibrium constants and their varia¬ 
tions with the temperature and the density must be 
known. The properties of the gases at high densities 
also depart widely from those of the ideal gas. In 
thermodynamics these properties are treated most 
simply by considering the properties the gas would 
have if it were ideal, and applying corrections for 




Figure 11. Equilibrium among the carbon atoms of the powder gases has been demonstrated by radioactive tracer 
experiments. 


CONFIDENTIAL 


































36 


PROPERTIES OF POWDER GAS 


departures from this state. For this approach, an 
adequate equation of state is necessary. 

Outstanding contributions to this thermo¬ 
dynamic problem have been made by several investi- 
gators . 24 - 340 - 435 - 440 - 441,480 The treatment of the prob¬ 
lem given here is necessarily brief, and the reader is 
referred for details to an NDRC report 24 in which is 
given a full discussion with applications and tabulated 
functions of the various quantities needed in com¬ 
putation. 

Conventions 

The notation employed is that of Lewis and Ran¬ 
dall . 511 The standard state chosen for elements is zero 
energy at 0 K, and for gases, that of the ideal gas. It 
may be noted that energies of formation (constant 
volume) and not heats of formation are employed 
throughout. The energies are obtained from conven¬ 
tional heats of formation Q f by equation (7), in which 

— A Et° = Q/ T* AnRT (/) 

An is the net increase in the number of moles of gases 
for the unit reaction. Thus, for the formation of water 
according to equation (8) 

H 2 (g) + ^0 2 (g) = II 2 0(1). ( 8 ) 

An = - 1.5and Atf 2 98.i6° = ~ (68,318.1 - 1.5 X 1.9871 
X 298.16) = - 67,429.4 cal/mole. 

Approximations 

It is assumed that there is no energy or volume 
change of mixing of powder gas constituents. Powder 
gas is thus treated as a perfect solution, and its prop¬ 
erties are computed additively from those of the in¬ 
dividual gases. Powder itself is treated as a simple 
mixture. These assumptions greatly simplify the work 
of computation; moreover, the contributions due to 
interactions of unlike molecules, although they are 
not too well known, are small for gas mixtures, except 
perhaps at extremely high densities. Experience shows 
that at densities of gas such as prevail in guns, the 
assumption of simple additivity is fully justified . 113 

2 4 2 The Equation of State 

for Powder Gas 

Actual gases at high pressures deviate greatly in 
their P-V-T relationships from the ideal gas. For ap¬ 
plication to the problem of internal ballistics, a modi¬ 


fied van der Waals’ equation in the form proposed by 
Abel, shown in equation (9), has been usually em¬ 
ployed. 

P(1 - 77 A) = nART. (9) 

Here A is the density of the gas, \/v, and rj is the co¬ 
volume which corrects for the apparent volume of the 
molecules. The equation neglects the van der Waals’ 
correction for the attractive forces between the mole¬ 
cules, as these may be expected to play only a minor 
role at the temperature of powder gas. 

For exact thermodynamic studies of powder gas 
such an equation does not reproduce the P-V-T rela¬ 
tions with sufficient accuracy. Of the numerous equa¬ 
tions of state proposed, equation ( 10 ), which is a van 
der Waals’ equation expanded in increasing powers of 
1/V, was selected for Division 1 ’s thermochemical cal¬ 
culations . 24 

PV h - a/RT h( h \ 2 

RT + V ^S\VJ 

+ 0.2869(A) 3 + 0.192s(Ay. (10) 

The quantities a and b are the van der Waals con¬ 
stants. The coefficients of the second and third power 
terms in (b/V) were calculated 5341535 - 536 - 537 to account 
for triple and higher collisions among the molecules. 
The fourth power term was added 454 to make the equa¬ 
tion merge smoothly into the limiting form for high 
densities, given by equation ( 11 ), 

If =F7^+[ 1 -°- 69C2 (v) 1 T < n > 

in which the coefficient 0.6962 is calculated for face- 
centered close packing of rigid spheres. For a body- 
centered lattice, the value would be 0.7163, and for 
simple cubic packing, 0.7816. 

The van der Waals a appears in these equations in 
the fraction a/VRT. Its effect on the pressure is small 
when T is high, and is ignored in ballistic applica¬ 
tions. High temperature values of b have been eval¬ 
uated by statistical mechanics ; 455 - 456 - 501 - 506 they are not 
strictly independent of T, but because the tempera¬ 
ture coefficient is small, they are assumed constant 
and equal to their values at 3600 K in the ballistic 
equation of state. Their values differ somewhat from 
the conventional van der Waals b’s derived from the 
properties of the critical state. Such values are listed 
in Table 2 for powder gas constituents. 

The covolume 17 in equation (9) has been evalu¬ 
ated 173 by analysis of the data 440 on uncooled pressures 


CONFIDENTIAL 





THERMODYNAMICS OF POWDER GAS 


37 


for the N(l) powder referred to in Section 2.4.5 and 
Table 4. A value of (1.074 ± 0.021) cc per g was recom¬ 
mended. 


Table 2. Van der Waals constants for constituents of 
powder gas. 


Constituent 

a 

Conventional* 

/ cm 3 V 
10 6 atml — 7- ) 
\mole/ 

b 

Conventional* 

cm 3 

mole 

High tempf 
cm 3 
mole 

C0 2 

3.60 

42.8 

63.0 

CO 

1.486 

39.9 

33.0 

h 2 

0.245 

26.6 

14.0 

n 2 

1.346 

38.5 

34.0 

h 2 o 

5.47 

30.5 

10.0 

nh 3 

4.17 

37.1 

15.2 

ch 4 

2.26 

42.8 

37.0 

NO 

1.341 

27.9 

21.2 

N 2 0 

3.79 

44.1 

63.9 

0 2 

1.361 

31.8 

30.5 


*Derived from critical constants. 
tCalculated by means of statistical mechanics. 


where, for a given gas, the energy of gas imperfection 
is defined by equation (15), 

E (gas imp.) = E (actual) — E (ideal). (15) 

For a mixture of gases, the energy of gas imperfection 
(in calories per gram of powder) is given by 

E (gas irnp.) mix = £n* E (gas imp.)*. (16) 

i 

The contribution to the energy of the gas from this 
cause rarely exceeds 10 cal/g at powder gas temper¬ 
atures. 

The Pressure of Powder Gas 

Assume the covolumes to be additive in a mixture 
of gases ( bmix/V = b'A) and define further a Active 
density according to equation (17), 

A' = 22.98mA, (17) 


For a mixture of gases, the gas imperfection correc¬ 
tion may be computed additively by the introduction 
of equation (12 ; or (13), 


bmix 

b' 


i 

= E^ibi, 


( 12 ) 

(13) 


where z* and m are the mole fraction and the number 
of moles of gas i in the mixture for 1 g of powder, 
respectively; b' is defined as the covolume. 

Important applications of the equation of state 
occur in the exact calculation of the pressure, com¬ 
position, and energy of powder gas. The correction for 
the energy of gas imperfection is small and may be 
neglected in the simplified calculation of the adiabatic 
flame temperature. The corrections for the pressure 
and the composition, on the other hand, are large, 
and neglecting them would lead to serious errors. 


The Change of Internal Energy with Density 
for Gases 

The internal energy of actual gases, unlike that of 
an ideal gas, varies with the volume (or the density). 
The contribution is relatively small, and arises almost 
exclusively from the temperature coefficient of 
b — a/RT. By statistical mechanical considerations, 
it has been evaluated for constituents of powder gas 
by equations of the type of equation (14), 

E (gas imp.)* = A'(A* - £* log T — CT l ) i (14) 


where n is the total number of moles of gas per gram 
of powder. Then the equation of state (10) can be 
solved for the pressure in atmospheres in accordance 
with equation (18), 

P = S.571A'T[1 + D(P,T)] (18) 

in which the numerical coefficient 3.571 is the value 
of R/ 22.98 in atm/cc/degree/mole and the function 
D(P,T), which is the gas imperfection correction, 
stands for the sum of the terms on the right-hand 
side of equation (10) minus 1. 


Composition of Powder Gas as a Function of 
Temperature 


The composition of powder gas varies with the tem¬ 
perature and density as a result of shifts in the equi¬ 
libriums of the chemical reactions. For a given reac¬ 
tion, the equilibrium constant K° for an ideal gas is 
related to the standard free energy change A F° for 
the reaction by equation (19). 


R In K° = 


-A F\ 


A {Ft - El) _ A El 


• (19) 


The second expression on the right of this equation is 
the more convenient to use because tabulated free 
energy changes are usually referred to the energy of 
the standard reaction at 0 K. 

Because the gas imperfection correction is different 
for different gases, values of K as ordinarily defined in 
terms of partial pressure, concentration, or mole frac- 


CONFIDENTIAL 









38 


PROPERTIES OF POWDER GAS 


tion x, vary with the density of the gas. The partial 
pressure Pi of gas i, for example, is equal to XiP for an 
ideal gas, where P is the total pressure. For actual 
gases at high pressure this is no longer true, but in¬ 
stead we may write equation (20), 


7t = 


yiPi 
XiP 


XiP’ 


( 20 ) 


in which y i} the activity coefficient for gas i, and /*, 
the fugacity of gas i, contain the correction for the 
gas imperfection, as expressed by equation (21), 

- -/ ¥§),"■ < 2i> 

The indicated integral may be evaluated from the 
equation of state; values of In y thus derived have 
been tabulated 24 for values of A' up to 0.5. 

The equilibrium constants may be written in terms 
of the fugacities = yiPi and related to K v , the 
equilibrium constant in terms of partial pressures, in 
accordance with equation (22), 

K f = KpKy = K p °. (22) 

K f then retains, at all pressures, the values of K p 
characteristic of the ideal gas. The correction for gas 
imperfection, contained in K y is conveniently intro¬ 
duced by equation (23). 

K p = KG = ~W~' W 

I\y A-y 

Values of G for the various reactions that may occur 
in powder gas as calculated from the values of y < ob¬ 
tained by equation (21) have been tabulated. 24 

When it is desired to evaluate the composition of a 
gaseous mixture at a given temperature T and den¬ 
sity A, it is convenient to deal with K n , the equilibri¬ 
um constant in terms of the relative numbers of moles 
of a constituent per gram of powder rather than with 
K p . The two constants are related by 



2 4 3 The Adiabatic Flame Temperature 

The Energy Released in the Burning of 
Powder 

The energy released per gram of powder is the 
energy given up to the surroundings when 1 gram of 
powder at 15 C is burned to form the (gaseous) prod¬ 
ucts at T K and density A g/cm 3 , as shown in equa¬ 
tion 24. f 


F rel (T°K,A) = E (powder, 15 C) 

— E° (elements, 15 C) 
+ (E°is C — E 0 °) elements 

— [E° (products 0 K) 

— E° (elements OK)] 

— ( Et° — E 0 °) products 
— E (gas imp.) 


E( 1) 
E(2) 

E{ 3) 
E( 4) 

m 


(24) 


The terms E( 1) and E{2 ) together represent the 
energy of formation of powder at 15 C from the ele¬ 
ments at 0 K; E( 3) is the negative of the energy of 
formation of the products from the elements, both at 
0 K; E( 4) is the negative of the energy required to 
heat the products in their standard states from 0 K to 
T K; and E{ 5), which is the negative of the energy 
due to gas imperfection, represents the energy neces¬ 
sary to convert the products from the standard state 
of ideal gas to the actual state. E{ 1) and E(2) are com¬ 
puted additively from the energies of formation of 
powder constituents referred to the elements at 0 K. 
E( 3), E{ 4), and E(S) are similarly computed additively 
from the corresponding energies for the gaseous prod¬ 
ucts. For this purpose, knowledge of the composition 
of powder gas at T Kand Ag/cm 3 is obtained from the 
values of KG, as outlined in the last paragraph of 
Section 2.4.2, and from the composition of the pow¬ 
der in terms of moles of its elements per gram of 
powder. 


Simplified Calculation of the Adiabatic Flame 
Temperature 

It will be noted that the highest temperature that 
can be attained adiabatically by the products is the 
temperature when the energy released is zero. This 
temperature is called the adiabatic flame temperature 
T 0 . It is an important constant for each powder. Its 


where a is the sum of the number of moles of the prod¬ 
ucts less the number of moles of reactants in the 
chemical equation. 


f In NDRC Report A-116 24 (p 32) the quantity represented 
here by the sum of E(l) and E( 2) is called the negative of the 
energy of formation of powder from the elements at OK; and 
the term E( 3) is designated as —E 0 °. 


CONFIDENTIAL 






THERMODYNAMICS OF POWDER GAS 


39 


evaluation by the general method just outlined is ex¬ 
ceedingly laborious, and a shorter method is desirable. 
It was found 23 that, since the energies are all com¬ 
puted additively, a considerable simplification could 
be achieved if a suitable compensation could be made 
for changes in the composition of the powder gas. It 
was noted that, while the gas composition varies con¬ 
siderably at a given temperature for different densi¬ 
ties, the energy released and the specific heat at con¬ 
stant volume are not greatly influenced by the den¬ 
sity except at very high temperatures, when account 
must be taken of energy-rich minor constituents. For 
the large group of powders with T 0 not greater than 
3000 K, equation (25) was proposed. 

To = 2500 (25) 

In it yi is the weight fraction of constituent i of the 
powder. 

The numerator of the fraction in equation (25) rep¬ 
resents the energy released at 2500 K, and the de¬ 
nominator the average value of C v over the range of 
temperatures 2500 ± 500 K, both computed additively 
from the composition of the powder. The energy re¬ 
leased at 2500 K and the C v were evaluated, first for 
the case of no C0 2 in the powder gas and second, for 


the case of no H 2 0. It was noted that typical powder 
gas at the temperatures and densities prevalent in 
guns contains 77 moles H 2 0 for every 23 moles of C0 2 . 
The mean energy released at 2500 K was therefore 
taken to be made up of the two extreme values in the 
ratio of 77:23. C v has a slightly larger value than 
would be expected on this basis, and a ratio of 1:1 of 
the two extreme values was found to give more nearly 
representative results. 

At temperatures above 3000 K it is necessary to cor¬ 
rect the energy released for the effect of dissociative 
equilibriums; when such a correction is made, equa¬ 
tion (26) results. 

To = 3000 - 60464 + 6046 (A 2 + B) H . (26) 
In this equation A and B have the values given by 
equations (27) and (28), respectively. 

A = T.ViCvi + 0.01185. (27) 

B = 0.0003308 (ZViEi - 500^2/,C,,i). (28) 

The values of To obtained by this approximate 
method usually agree very closely with those obtained 
by the accurate method. Table 3 gives values of the 
additive constants of powder constituents, C„», £’,• 
(2500 K), ui and rji for reference. 


Table 3. Molar additive constants of powder constituents per gram of powder for computation of adiabatic flame 
temperature.* 


Constituent i 

c vi 

Ei 

Ui 

Vi 

Nitrocellulose f 

0.3421 

274.6 

0.03920 

27.56 


(0.006Y) 

(-142Y) 

(—0.00218Y) 

(+1.00Y) 

Nitroglycerine 

0.3439 

951.9 

0.03083 

22.78 

Diphenylamine 

.3475 

-3009.7 

.10637 

65.44 

Dibutylphthalate 

.4261 

-2694.7 

.09700 

56.97 

Dinitrotoiuene 

.3213 

- 708.4 

.06040 

40.44 

Water 

.6507 

-1567.6 

.05551 

24.60 

Acetylene 

.3755 

-1374.0 

.11523 

69.30 

Acetone 

.5107 

-2842.5 

.10331 

57.22 

Ethyl alcohol 

.6085 

-2784.8 

.10854 

56.35 

Ethyl ether 

.5980 

-3073.7 

.12142 

64.06 

Centralite 

.3909 

-2873.7 

.10444 

62.60 

Nitroguanidine 

.3711 

- 60.5 

.04804 

31.77 

Cyclonite (RDX) 

.3415 

622.3 

.0405 

28.5 

PETN 

.3485 

724.1 

.0348 

24.9 

Vaseline 

.5983 

-4175.1 

.142 

65.6 

Diamylphthalate 

.4408 

-2809 

.1013 

58.8 

Trinitrotoluene 

.3035 

- 110.1 

.0484 

34.3 

Triacetin 

.4191 

-1973 

.07331 

43.77 

Graphite 

.1349 

-3223.8 

.08326 

60.53 

NH4NO3 

.4424 

405.1 

.03748 

22.83 

kno 3 

.2158 

25 

.00989 

23.61 

Ba(N0 3 ) 2 

.1574 

131 

.00765 

15.34 

k 2 so 4 

.1250 

- 860 

.00574 

9.36 


* Taken from NDRC Reports A-101 23 and A-142. 26 The values of Ei for nitrocellulose, dinitrotoiuene, centralite, and nitroguanidine have been recalcu¬ 
lated by W. S. Benedict from thermal data furnished by F. D. Rossinil 561 > 562 Values for ether have also been added. The values of ni have been checked 
and modified when necessary, 
t Y = 13.15 - %N. 


CONFIDENTIAL 








40 


PROPERTIES OF POWDER GAS 


2 4 4 Other Thermodynamic 

Properties of Powder Gas 

For completeness it is desirable to evaluate the 
enthalpy, the entropy, and the specific heats of pow¬ 
der gas. 

The Enthalpy 

The enthalpy (also called the "heat content” or 
“total heat”) H is defined by equation (29). 

H = E + PV (29) 

When referred to its value at 0 K, the enthalpy rep¬ 
resents the heat that must be added to raise the tem¬ 
perature of the gaseous products from 0 K to T K, the 
products being maintained in thermal and chemical 
equilibrium throughout, at the equilibrium pressure 
P( T) corresponding to the given density A g per cu 
cm. The E term in the definition is — [P rel (0 K) 
~-^rei(TK)], taken negative because the energy re¬ 
leased is given up by the system. The PV term is zero 
at 0 K, since the equilibrium pressure at this temper¬ 
ature is zero at all densities. Accordingly the enthalpy 
at 0 K, Ho equals — P rel (0 k)> and the enthalpy H 
at any temperature T referred to its value at 0 K 
is given by equation (30), 

H = (. Ht — Ho) = E re \ (o k) — E r «\(TK) 

+ 0.02421 UPV, (30) 

which may be written as equation (31) 

H — Hi — E rei(TK) + 0.024214 — (31) 

in order to introduce the symbols used in the Division 
1 system of ballistics (Section 3.1), H being expressed 
in calories per gram and P in atmospheres. 

The Specific Heats 

The ratio of specific heats, y = C p /C v , rather than 
the separate values of C p or C v , finds extensive use in 
the calculations of interior ballistics. It will be evi¬ 
dent that, when the energies released and the enthal¬ 
pies are tabulated for different values of T and A, C v 
and C p may be obtained from equations (31) and (32). 

Cv = - A Cal /g- (32) 

Cp = ( ^ cal/g. (33) 


The ratio of the specific heats for the gaseous prod¬ 
ucts in the state of the ideal gas ( 7 0 ) is given by 
equation (34), in which C v ° in calories per mole may be 
obtained from tabular differences of Ui(E T Q — Po°)i at 
two different temperatures T, summed up for the 
separate powder gas constituents i. 


7° = 1 + 


nR 

~C? 


(34) 


= 1 - 


(35) 


At high densities and pressures, where the energy of 
gas imperfection may become significant, an accurate 
value of 7 may be obtained from equation ( 35 ), 

0.0242147^1 l(dA\ ( dP\ 

C p A 2 / \W/,\dA/\ 

in which C p is given in cal/g/degree K, P is in atmos¬ 
pheres, and the other symbols have their usual sig¬ 
nificance. The derivatives are obtained from graphs of 
(A vs P T and of (P/A vs T) p by use of the relation ex¬ 
pressed by equation (36). 


- 2 3+l»4ML - < 36 > 

An effective ratio of the specific heats ( 7 e ff) may be 
obtained on the assumption of equation ( 37 ) combined 
with equation (38), 


(C p - C t ) = nR (37) 

= 1 + (38) 

-ftreli 0 

where (F) is the “force” or impetus, nRTo. It may be 
noted here that the influence of pressure on the spe¬ 
cific heats or their ratio is small, and arises from the 
small temperature coefficient of the energy of gas im¬ 
perfection. 


The Entropy 

The entropy, aside from its intrinsic value, has use 
in the construction of Mollier diagrams. Equation (39) 
expresses the entropy for the mixture represented by 
powder gas. 

_ E t ° —Eo° E (gas imp.) F T ° — E 0 ° 

^ rp "l rji rjn 

— nP[ln (3.571 7 7 A / ) + < In Xi t 4> — 1]. (39) 

1 

In using this equation (E T ° — Po°) and (F T ° — E 0 °) are 
obtained additively on multiplying the number of 
moles Ui and the respective energies for the different 
gases, and 4> is given by equation (40). 

4> = b'A + 0.3125(6'A) 2 + 0.09563(6'A) 3 

+ 0.0482(6'A) 4 . (40) 


CONFIDENTIAL 





THERMODYNAMICS OF POWDER GAS 


41 


2 4 5 Applications of Thermodynamics 
to Powder Gas 

Thermodynamics has been applied 24 to the calcula¬ 
tion of properties of powder gas from a number of differ¬ 
ent propellants. Space does not permit considering 
these in detail.Types of results obtained are illustrated 
in Figure 12, giving the energy released for a nitrocellu¬ 
lose powder [designated 440 “N(l)”], similar to theU.S. 
Navy Pyro powder, an FNH powder, lot No. 1358 
(designated “D”) and a cordite [designated 440 - 538 
“C(l)”] containing 28.7% nitroglycerine. In each case 
the E nX is given for two densities, as indicated in the 
figure. At the lower temperatures, E nl is not a sensi¬ 
tive function of the density, but near 3000 K and 
higher the dissociation of the main products, which is 
influenced by the density, becomes appreciable, and 
the energy released reflects this. The adiabatic flame 
temperatures To for these powders are the tempera¬ 
tures at E IeX = 0. They are for A = 0.25, 2915 K for 
powder N(l), 2575 K for powder D, and 3735 K for 
powder C (1). The values calculated by a short meth od 23 
are 2899, 2577, and 3727 K. Calculated values of the 
thermodynamic properties of British powders have 
been tabulated. 9 

2 - 4,6 Comparison of Calculated Thermo¬ 
dynamic Properties with Experimental Ones 

Pressures 

The pressures calculated thermodynamically are 
the uncooled pressures—that is, no allowance is made 
for heat loss to the walls. Measured pressures are 
lower. A method 191 of correcting observed pressures 
for heat loss has been derived at Aberdeen Proving 
Ground. A comparison of the calculated pressures at 
To with corrected observed pressures for powder gas 
from powder N(l) are given in Table 4. 

Heat of Explosion 

The heat of explosion, per gram of powder, as ob¬ 
served in a closed chamber calorimeter is equahto the 
energy released at the calorimeter temperature (water 
liquid). It is of interest to determine the influence of 
the gas composition on the energy released, and com¬ 
pare the result with a measured heat of explosion. The 
powder selected for this is powder N(l), for which a 
calorimetric value 440 at a density of loading 0.1 is 936 
cal per g (water liquid, no gas analysis is given). In 


Table 4. Comparison of calculated uncooled pressures 
for N(l) powder gas with experimental values of Crow 
and Grimshaw as corrected by Kent and Vinti. 191 


Density 

A 

(g/cm 3 ) 

Po(calc) 

(kg/cm 2 ) 

P o(expt) 
Large chamber 
(kg/cm 2 ) 

Po(expt) 

Small chamber 
(kg/cm 2 ) 

0.2479 

3368 

3408 

3391 

.2211 

2911 

2946 

2934 

.2049 

2647 

2690 

2729 

.1819 

2288 

2288 

2336 

.1546 

1883 

1907 

1896 

.1344 

1597 

1614 

1618 

.1066 

1225 

1232 

1209 

.0755 

835 

826 

823 

.0585 

633 

632 

636 

.0420 

446 

435 

442 

.0263 

274 

270 

266 


making the calculations we have adopted 300 K as the 
calorimeter temperature and assumed the water gas 
equilibrium to freeze at a series of temperatures Tj 
from 300 to 3000 K. In view of the quenching results 
discussed in Section 2.3, the values for T = 2000 K 
and above obviously have no physical significance 
and are included for another purpose to be made 
evident shortly. The results, considering water gas 
constituents and nitrogen only at densities A' = 0, 
0.1, 0.25, and 0.4 are given in Table 5, for water gas- 



TEMPERATURE IN DEGREES K 


Figure 12. Energy released versus temperature for 
three different powders. (This figure is based on data in 
NDRC Report A-116.) 


CONFIDENTIAL 
























42 


PROPERTIES OF POWDER GAS 


Table 5. Energy released at 300 K for Crow and 
Grimshaw’s 440 N(l) powder, calculated for the indicated 
values of A', and T g ', water-gas constituents only. 
(A = 1.05943A' for this powder.) 


77/A' 

300 

1000 

1500 

2000 

2500 

3000 

0 

(911.4)* 

911.4f 

(858.2) 

911.8 

(830.7) 

912.0 

(819.7) 

912.2 

(814.3) 

912.2 

(811.4) 

912.2 

0.1 

(922.1)* 

922.lt 

(863.7) 

921.6 

(835.9) 

921.2 

(825.0) 

921.1 

(819.7) 

921.1 

(816.9) 

921.0 

0.25 

(938.3)* 
938.3 f 

(864.7) 

935.5 

(837.7) 

934.4 

(827.7) 

934.1 

(823.0) 

933.8 

(820.6) 

933.8 

0.40 

(954.4)* 
954.4 f 

(854.8) 

948.0 

(832.1) 

946.5 

(824.6) 

946.0 

(821.3) 

945.8 

(819.6) 

945.7 


* E rei 300 k water gaseous, 
t E re i 300 K water liquid. 


eous (in parentheses) and for water liquid. The feature 
to be noted is the negligible effect of Tj and hence, of 
the shift of the water gas equilibrium on the energy 
released (water liquid). 

The value of the heat of explosion calculated for 
A =0.10 (A' = 0.0944 with n = 0.041075 mole/g) for 
T 0 ' = 1500 K is 92.07 cal/g.The methane formed and 
its influence on the properties of the gas at 1500 K 
had been calculated 24 previously. The presence of 
methane, because of the large amount of heat evolved 
in its formation, increases the energy released. If its 
equilibrium quantity at 1500 K were formed, and the 
equilibrium froze at this temperature, E Tei 30 o k at 
A = 0.10 would be increased by 18 cal/g, making a 
total of 939 cal/g, in comparison with the measured 
heat of explosion of 936 cal/g. The agreement is excel¬ 
lent. Actually the methane equilibrium tends to 
freeze-in at a Tj of about 1800 K,which would mean 
a somewhat lower than equilibrium concentration 
at 1500 K, and a lower contribution to the energy 
released. 

The use of a rational heat of explosion, defined as 
the amount of heat liberated per gram when the pow¬ 
der gas cools from the adiabatic flame temperature to 
room temperature without a change in composition 
has been advocated. 539 The rational heat is identical 
with the E reX at the calorimeter temperature, (water 
gaseous), with Tj = T 0 , A = A 0 (density of loading), 
provided all the powder burns. For N(l) powder gas 
at T 0 f = T 0 = 2915 K and A = 0.10 it is 817.5 cal/g, 
as interpolated from Table 5, values in parentheses 
(water gaseous). 

Among the many applications of thermodynamics 
to powder gas that might be mentioned, an out¬ 
standing example is given by an interpretation of the 
chemical thermodynamics of gun erosion. 60 In this 


work the powder gas is assumed to be quenched at a 
series of temperatures T a ' as defined earlier in this sec¬ 
tion, and the free energy of the reaction of the 
quenched gas with the surface of the bore at its sur¬ 
face temperature T s is calculated. The work is pre¬ 
sented in detail in Sections 12.2.4 and 13.3.3, together 
with a slightly different approach 138 to the same 
problem. 

2 5 POWDER GAS RADIATION AND 
TEMPERATURE 

251 Temperature Determinations 
from Radiation Measurements 

Thermodynamic calculations for powder gas yield, 
as is shown in Section 2.4, a value for the uncooled 
pressure Po, and the adiabatic flame temperature To. 
In interior ballistic calculations certain simplifying 
assumptions based on experience are made concern¬ 
ing the burning rate, heat, and frictional loss, and 
these with a simplified equation of state for the gas 
enable the course of the pressure and temperature to 
be computed as a function of the time or travel as 
described in Chapter 3. On the experimental side, the 
advent of the piezoelectric and the electrical-resistance 
strain gauges make possible a measurement of 
the pressure as a function of time as described in 
Chapter 4. 

Until recently, however, there has been no reliable 
measurement of the temperature for powder gas. In 
the past, attempts to measure the temperature have 
been made with thermocouples, and by inclusion in 
the propellant of metals in thin sheets or in pulver¬ 
ized form. These methods have failed to supply ade¬ 
quate information for reasons which need not be dis¬ 
cussed here. More recently studies were made in 
England of the spectral distribution of the emitted 
radiation, generally at low density of loading. 352 - 358 - 393 
These have shown a continuous background radia¬ 
tion together with lines of Na, K, and CaO bands. 

In an investigation 44 at the Geophysical Laboratory 
for Division 1 experimental methods were developed 
for precise measurements of the radiation from burn¬ 
ing propellants at densities of loading corresponding 
to those in guns. The stumbling-block in making 
radiation measurements at high densities has been 
the lack of a strong enough window. Successful use 
was made at pressures up to 50,000 psi of fused quartz 
windows of the Poulter type 540 in which a plane 
polished window surface is brought in contact with a 


CONFIDENTIAL 








POWDER GAS RADIATION AND TEMPERATURE 


43 


plane polished surface of a hardened steel supporting 
plug. 

The first results of this investigation showed that 
at a sufficiently high density of loading (0.03 mini¬ 
mum, with a path length of 1.75 cm) the radiation is 
not only continuous, but has a spectral distribution 
closely resembling that of a black body, 8 and that 
hence, a powder gas temperature can be determined 
from measurements of the radiation intensity. 

This principle was then applied to the develop¬ 
ment of the photoelectric pyrometer described in Sec¬ 
tion 2.5.4. With it studies have been made of the 
temperature of powder gas in small closed chambers, 
in a jet propulsion motor, and in the 3-in. gun at 
Carderock described in Section 4.2.2. The results are 
presented briefly in Section 2.5.5. Preparations 109 ’ 110 
were made for similar measurements to be carried 
out during the firing of a 90-mm gun, M1A1 at Car¬ 
derock (Section 4.1). 

2 52 Spectral Characteristics 

of Powder Gas Radiation 

Spectra of radiation emitted by powder gas at very 
low densities of loading or at early times in firings at 
intermediate and high densities of loading in a closed 
chamber consist of discrete lines and bands superim¬ 
posed upon a continuous background. In specific ex¬ 
periments the discrete features observed had their 
origin in resonance transitions of potassium, sodium, 
calcium, (andcalcium oxide), barium, iron, chromium, 
nickel, and vanadium, in the order of their prominence. 
A hydroxyl (OH) band at 3064 A was observed in 
absorption, and a faint band near 8150 A may be a 
vibrational bard of water. The chromium, nickel, and 
vanadium, together with some of the iron, probably 
came from the nichrome wire igniter; the other sub¬ 
stances were present in powders either as impurities, 
or as a deliberate addition to make the powders 
“flashless.” 

The intensity of the continuous radiation increases 
with the gas density and length of path; as the density 
increases some of the stronger discrete lines may ap¬ 
pear in reversal. The sole origin of the lines and bands 
in resonance transitions, and the intensity of the con¬ 
tinuum are evidence that the spectra are thermally 
excited. 


g The black body characteristics of powder gas radiation 
were earlier studied by Col. Libesart 487 who, however, does not 
appear to have carried the matter further. 


None of the major constituents of powder gas emit 
strongly enough in the visible and the near infrared, 
in comparison with the inorganic impurities, to make 
a detectable contribution to the spectrum, with the 
possible exception of water vapor. Among the free 
radicals, only OH, CN, CH, and NH might be de¬ 
tectable. 

The continuous spectra may have their origin in 
(1) recombination of free radicals which would not 
lead to black body distribution of intensities, or (2) 
glowing “soot” or inorganic particles. The latter 
alternative is much the more probable. It has been 
shown that particles of diameter 1 /jl or less assume 
the temperature of the gas in times of the order of 
0.1 msec. The radiation in such a case would have a 
distribution closely approximating that of a black 
body at the temperature of the gas. 

A characteristic spectrum obtained with a moving 
film spectrograph at a density of loading 0.082 of 
Hercules No. 2 pistol powder is shown in Figure 13. 
Microphotometer traces of this and similar spectra, 
together with comparison spectra of the tungsten fila¬ 
ment heated at known color temperatures, are repro¬ 
duced in Figure 14, where the legend explains the vari¬ 
ous experimental conditions. It will be noted that as 
time, and hence the gas density, increases during a fir¬ 
ing, the relative prominence of the sharp peaks cor¬ 
responding to line structure decreases until all the 
tracings assume the same general shape. The coinci¬ 
dence of the tungsten strip and the powder gas spec¬ 
tra is a strong indication that, as the density increases, 
the spectral distribution of the continuous radiation 
approaches that of a black body, at least for the wave¬ 
length range 4000 to 6400 A observed in the experi¬ 
ments. 

Measurements of the emissivity and absorptivity 
of powder gas at various densities also showed that 
unit emissivity is rapidly approached at values of the 
density times the length of path about 0.06 g/cm 2 . 
The actual value of the density at which a stated emis¬ 
sivity e is reached depends on the mass emissivity k 
characterizing the gas from a particular powder, as 
given in equation (41), 

In (1 — e) = — kAl (41) 

where A is the gas density and l the length of path. 

Values of k found for NH andFNH-M2 powder gas, 
for example, were 34 and 120 cm 2 /g. In a gun at a time 
when A = 0.2 g/cm 3 , e = 0.9 would be expected to be 
reached with such powders in thicknesses of gas 0.34 
and 0.095 cm, and e = 0.99, in 0.68 and 0.19 cm, re- 


CONFIDENTIAL 




44 


PROPERTIES OF POWDER GAS 



Figure 13. Spectrum of burning powder photographed with a moving-film spectrograph, length of path 1.75 cm. Her¬ 
cules No. 2 pistol powder, loading density 0.082, a comparison spectrum of the mercury arc appears superposed on the 
shot. (This figure has appeared as Figure 15 in NDRC Report A-252.) 



Figure 14. Four sets of microphotometer traces from the spectra obtained with the rotating drum spectrograph, princi¬ 
pally from Figure 13, together with comparison spectra of a tungsten strip-filament at known color temperatures. All the 
films were developed at the same time. (This figure has appeared as Figure 16 in NDRC Report A-252.) 


Set 

Trace 

Source 

Time (msec) 

A 

a 

Shot, Powder (2), Ao =0.082 

0.1 

A 

b 

Shot, Powder (2), = .082 

.25 

A 

c 

Shot, Powder (2), = .082 

.5 

A 

d 

Shot, Powder (2), = .082 

.75 (max) 

B 

a 

Shot, Powder (2), Ao =0.082 

0.3 

B 

b 

Shot, Powder (2\ = .082 

1.3 

B 

c 

Tungsten, T c =3120 K 


C 

a 

Shot, Powder (2), Ao =0.082 

0.60 

C 

b 

Shot, Powder (2\ = .082 

1.00 

c 

c 

Tungsten, T c =3240 K 


D 

a 

Shot, Powder (2), Ao =0.082 

0.45 

D 

b 

Shot, Powder (2), = .082 

1.05 

D 

c 

Shot, Powder (19), = .179 

(max) 


CONFIDENTIAL 

































POWDER GAS RADIATION AND TEMPERATURE 


45 


spectively. Gases from powders containing large 
amounts of inorganic constituents tend to be char¬ 
acterized by high values of k. 

2 - 53 Characteristics of Radiation 
Black Body Radiation 

The intensity of radiation emitted by a body in 
thermal and radiative equilibrium J(X, T ), at the abso¬ 
lute temperature T and at the wavelength X is accur¬ 
ately expressed by the well known Planck distribution 
law; for temperatures below 4000 K and wavelengths 
in the ultraviolet, visible, and near infrared, it is ex¬ 
pressed without appreciable error (less than 5 degrees) 
by the simpler Wien equation, equation (42\ in which 
Ci = 2irhc 2 = 3.740X10 -5 erg cm 2 /sec -1 and = hc/k 
= 1.436 cm/deg. 

J(A, T) = ci\ -5 exp - - C2 , (42) 

At 

Equation (42) has been confirmed by prior experi¬ 
ments for wavelengths in the visible and ultraviolet 
regions, even at the highest temperatures. 

Radiation from Actual Bodies 

As a consequence of reflection and the possible fail¬ 
ure of Lambert’s cosine law for solids or liquids, or of 
low mass emissivity k for gases, real bodies may ra¬ 
diate less than a black body. The ratio of the surface 
brightness B(A,T) to that of a black body R 0 (X,7 7 ) 
= J(A,T) is given by the spectral emissivity e(X,T) 
which in general is less than unity and varies with 
both the temperature and the wavelength. The spec¬ 
tral distribution of brightness for such bodies there¬ 
fore differs from that of a black body. 

For any one selected wavelength X* it is possible, 
however, to define an apparent brightness temperature 
Si , by Wien’s law such that a black body at Si has the 
same brightness as the real body at T, as shown in 
equation (43) 

B(A i} T) = CiXr 5 exp —(43) 

A i&i 

Si will in general be different for every X*. 

It is also possible to seek a second apparent tem¬ 
perature, the color temperature T c , such that a black 
body at T c has the same spectral distribution as the 
real body at T, as expressed by equation (44). 

= Cl \- 6 exp (44) 


Such a color match may not be possible for all wave¬ 
lengths, but it can be made for two selected ones, such 
as 6650 and 4670 A customarily used in pyrometry. 
The apparent temperature T ci j thus defined is usually 
termed the two-color or the red-blue color temperature. 
The three temperatures, Si, T cij and T, the spectral 
emissivity e* and the two-color emissivity e cij are con¬ 
nected by the relation given in equation (45). 


B(Ai, T) = CiXr 5 exp — 

= Cl €iXr 6 exp yy 
= Cie cij \r 5 exp - S 2 - . (45) 

Ai 1 cij 


where Si is defined for any one X;, T cij for any two 
Xi and Aj and T for all values of X. 

In particular, the ratio of the brightness at two 
wavelengths X r and A b for a body at the temperature T 
is given by equation (46). 


In 


B(\r, T) 
B(A b ,T) 


5 In y + %--( i ~ y)- ( 46 ) 

Ar -I crb \ Aft A r / 


From this equation it is evident that T c = T if e r /e b 
= 1 even though e r and e b separately are not unity. 
Since the ratio e r /e b for two fixed wavelengths A r and Ah 
is only a weak function of the temperature, it may be 
assumed constant without an appreciable error; equa¬ 
tion (46) may then be expressed by the simpler equa¬ 
tion (47), 


In _ Af _j_ 

B{A b ,T) ^ T 


A + 


czB 


(47) 


which is linear in reciprocal T or T c . 

As has alread}^ been stated, the emissivities of pow¬ 
der gas approximate unity in a sufficiently thick layer 
of gas even at quite low densities; when this is true, 
the brightness, color, and the true temperature should 
coincide. For substances having relatively low emis¬ 
sivities, however, the differences may be quite large, 
and dependent on wavelengths. Thus, for tungsten, 
which is of interest because it is used in making cali¬ 
brations, it is found at T = 2500 K, with X r = 6650 A, 
Xb = 4670 A and e r = 0.4250, = 0.4620, that Sr 

= 2275 K, S b = 2352 K, e c = 0.349 and T c = 2558 K. 
If, on the other hand, A r = 9000 A, A b = 4500 A with 
e r = 0.3610,6b = 0.4655, then S r = 2156K, S b = 2359K, 
e c = 0.2800, and T c = 2604 K. 


CONFIDENTIAL 



46 


PROPERTIES OF POWDER GAS 


2,5 4 Photoelectric Pyrometer 

for Powder Gas 

Design Principles 

In the burning of powder in closed chambers and 
guns the total time in which a quantitative measure¬ 
ment of the radiation must be secured is of the order 
of a few milliseconds, and it is desirable to have a fre¬ 
quency response up to 10 5 c. Among the possible 
methods of securing such a measurement, the most 
advantageous appears to be the use of vacuum-type 
photocells to observe the radiation, combined with os¬ 
cillographic recording of the photocell response, as a 
function of time. In photocells the response is in¬ 
stantaneous, directly proportional to the intensity of 
radiation over a wide range, and moreover, cells are 
available with spectral sentivities such that different 
spectral regions may be explored. 



Figure 15. A two-color photoelectric pyrometer for 
mounting on a 3-in. gun. (This figure has appeared as 
Figure 26 in NDRC Report A-323.) 


Measurements for calculating the brightness tem¬ 
perature Si [defined by equation (43) ] may be made 
with single cells; it is more desirable, however, to make 
measurements with two cells simultaneously for de¬ 
termining the color temperature T c [defined by equa¬ 
tion (44) ], inasmuch as many factors which assume 
undue importance in the single-cell measurements bal¬ 
ance out when two cells are used. This is especially 
true concerning the absorption of radiation by the 
window, the transmission through which varies with 
time owing to the deposition of smoke particles. 

The isolation of wavelengths or wavelength ranges 
transmitted to the photocells can be made by a mono¬ 
chromator or by filters. Because of the usual small 
aperture of monochromators the oscillograph deflec¬ 
tion is very small unless the photocell response is over¬ 
amplified ; for application to recoiling guns their use 
would be unworkable. The method was tested, how¬ 
ever, and for research purposes with instruments of 
high light-gathering power, is deserving of further 
trial, particularly with low-density firings when 
deviations from black-body conditions may be im¬ 
portant. 

For routine measurements wavelength ranges may 
be isolated by appropriate glass or gelatine filters. 
Suitable combinations for the two-color measurement 
are the RCA phototube 922 or 925 with Wratten filter 
87, giving an effective X of 9000 A and 8600 A respec¬ 
tively, and an RCA phototube 929 with filter 556 
effective X 4450 A; with filter 511, effective X 4200 A; 
or with filter 43, effective X 4800 A. Cell 929 can also 
be used without a filter when its effective X is 5000 A. 

Apparatus 

For photoelectric measurements by the two-color 
method, the radiation as it emerges from the window 
in the pressure vessel or gun must be divided so that 
separate beams may reach simultaneously the red- 
sensitive and blue-sensitive photocells. This may be 
done by a spectrograph (provided with two emergent 
slits) or better, by a semitransparent mirror. One of 
the “pyrometers” incorporating this feature and de¬ 
signed for mounting on a 3-in. gun is illustrated in 
Figure 15. (See Section 4.4.13.) 

In this installation the radiation from the window 
W passes through a condensing lens of suitable focal 
length, and is divided by the mirror so that part of the 
beam is reflected to the “red” cell, and part trans¬ 
mitted to the “blue” cell. Gelatine filters are wrapped 
around the phototubes. Since the pyrometer recoils 


CONFIDENTIAL 
































POWDER GAS RADIATION AND TEMPERATURE 


47 


with the gun, relative vibration of the photocell ele¬ 
ments may be objectionable. To minimize this, the 
aluminum box holding the cells is arranged to swivel 
around the retaining plug so it may be oriented for 
the axes of the cells to be parallel to the direction of 
motion. 



Figure 16. Oscillographic calibration record for the 
two-color photoelectric pyrometer, giving traces cor¬ 
responding to the stated color temperatures of a tung¬ 
sten strip-filament, at three amplifications in the ratio 
1:5:10. As the temperature increases, the blue/red ratio 
increases. The dot and dash curves are curves of con¬ 
stant emissivity for the two lower amplifications. (This 
figure has appeared as Figure 10 in NDRC Report 
A-252.) 

The output of the photocells is led first to a pre¬ 
amplifier giving fixed gains adjustable in steps, and 
then to a cathode ray oscillograph. The output of each 
photocell may be recorded separately as a function of 
time by a rotating drum camera or by using a recur¬ 
rent sweep and a still camera. For the two-color meas¬ 
urement, the outputs of the two cells are connected 
individually to the X- and Y-axes of the oscillograph; 
the spot then moves at an angle which is a measure of 
the ratio of the two amplified photocurrents, and 
thereby, becomes an index of the color temperature. 
The amplitude of the motion of the spot is similarly a 


measure of the emissivity of the gas; the absolute 
value of the emissivity may be determined from the 
ratio of the observed photocurrent to that which 
would be produced under like conditions by a black 
body at the same temperature. 

The calibration 44 of the instrument is carried out by 
comparison with a tungsten-strip lamp operated at 
known brightness temperatures, calibrated with refer¬ 
ence to an optical pyrometer. A typical calibration 



Figure 17. Results of a series of calibration records 
made by the two-color photoelectric pyrometer. The 
photocurrent is plotted versus the color temperature of 
the tungsten filament. Circles represent experimental 
points; the slopes of the R and B lines were calculated 
from the phototube sensitivities and filter transmissions 
in combination with the Wien formula and the known 
emissivity of tungsten. Vertical adjustment of the curves 
is made to fit the individual sensitivities of phototubes 
used. (This figure has appeared as Figure 11 in NDRC 
Report A-252.) 


record obtained with the two-color pyrometer is re¬ 
produced in Figure 16, and a plot of a series of calibra¬ 
tions is given in Figure 17, where the logarithms of the 
photocurrents of the red and the blue cells and their 
ratios are plotted against the reciprocals of the color 
temperature T c . 


CONFIDENTIAL 



















48 


PROPERTIES OF POWDER GAS 


2 5 5 Determinations of 

Powder Gas Temperatures 

Firings in Closed Vessels 

During the course of development of the photo¬ 
electric pyrometer, many firings with four typical 
propellants (NH, FNH-M2, single-base pistol, dou¬ 
ble-base pistol) were made at loading densities from 
below 0.01 to more than 0.25 in closed vessels pro¬ 
vided with windows. Only one series of six shots at 
A = 0.20 with NH powder, designed to test the re¬ 
producibility, will be discussed here brief y. 



Figure 18. Oscillographic temperature record in a 
closed chamber firing taken with a two-color photoelec¬ 
tric pyrometer. Arrows indicate the advance of time. 
Calibration traces at color temperatures of 249IK and 
2604 K are superposed. 

Figure 18 gives a two-color oscillograph record for 
the first shot in the series, with three superimposed 
calibration traces. The direction of motion of the 
oscillograph spot is indicated by the arrows. Timing 
dots on the trace give a correlation of events with 
pressure-time, and with red (or blue) radiation in¬ 
tensity-time traces taken simultaneously with other 
oscillographs. From measurements of the ordinates 
and abscissas at any point on the trace the photo¬ 
currents and their ratios are evaluated, and from 
these, with reference to the calibration chart, the 


temperatures. Once the temperatures are known, the 
apparent relative emissivities (e r )r or (e b )r can be 
calculated by comparing the photocurrent in firing 
with the photocurrent in calibration for the particu¬ 
lar value of T. 

The results for the six shots in the series are sum¬ 
marized in Figure 19 which gives the pressures and 
the emissivities, referred to their values at rupture 
(release of pressure), and the individual values of the 
two-color temperature T c for the various firings. The 
pressures appear to be reproducible to within 2 per 
cent of the average curve. The emissivities show a 
much greater variation, but in general they rise to 
the saturation value before the pressure rises to 20 
per cent of its final value. 

The general course of the temperature with time is 
shown to be similar for all the six shots: with indi¬ 
vidual variations, there is an initial peak followed by 
a drop, and then a more gradual rise to the final 
value. It will be noted that part of the time the 
emissivity exceeds unity. This appears to be common 
for e r while the powder is actively burning, and may 
be due to the same causes as the strong infrared 
emission observed in a separate investigation 160 al¬ 
ready referred to in Section 2.2.4. The individual 
variations in temperature may arise partly as the 
result of irregular ignition and burning of powder 
grains; more specifically they emphasize the inherent 
variations in the local temperature of gas in turbu¬ 
lent motion. When the gas is comparatively opaque 
(high mass emissivity k), the radiation which is emit¬ 
ted comes from a comparatively thin layer next to 
the window which may be cooler than the average, 
and whose temperature may be varying because of 
local pressure inequalities and other sources of imper¬ 
fect thermal equilibrium. 

Firings in 3-in. Gun 

As part of the comprehensive series of internal and 
external ballistics measurements with the 3-in. gun at 
Carderock, described in Section 4.2.2, measurements 
for determining the powder gas temperature were 
taken, first through holes into the chamber alone, and 
later, in two other positions, one 10 in. and the other 
51 in. ahead of the forcing cone (see Section 4.3.16). 
Firings were made with Navy “SPDN” and “SPD” 
powders (NH and pyro, respectively) at full charge, 
and at fractional charges down to 0.50 of full charge. 
The details are given in separate reports. 39 - 65 132 Here 
it is possible to reproduce only two examples of the 


CONFIDENTIAL 






6,JT 


POWDER GAS RADIATION AND TEMPERATURE 


49 


results obtained. Measurements were taken with the 
two-color pyrometers, and wiih single-color pyrom¬ 
eters ; the temperatures were determined as the two- 


color temperature T c or the single-color brightness 
temperatures T R or T B - 

Figure 20 gives a correlation of the temperature, 



TIME IN MSEC 

Figure 19. A diagram illustrating the reproducibility of results in six consecutive firings in a closed chamber, with NH 
powder. The absolute temperature T, the relative emissivity (e=et/or), and the pressure (ir = Pt/Pr ) are plotted as 
functions of time. The emissivities and pressures are expressed in terms of their values at rupture. Single curves are drawn 
to represent average e and tt for all shots; the T s are sketched individually up to the time of rupture; thereafter they 
are roughly parallel. (This figure has appeared as Figure 24 in NDRC Report A-252.) 


CONFIDENTIAL 


TEMPERATURE IN DEGREES K 
















































50 


PROPERTIES OF POWDER GAS 



87654321 0 


0.50 4 5 


0.25 40 


0.00 35 


9.75 30 


CO 

(L 

9.50 § 25 
uj 2 

§ ? 

UJ 

9.25 * 20 

CO 

cn 

UJ 

a: 

a. 

9.00 15 


8.75 10 


8.50 5 


8.25 0 


TIME TO EJECTION IN MSEC 

Figure 20. Powder gas temperature, pressure and emissivity curves for a 100 per cent round with SPDN (NH) powder 
in 3-in. gun, plotted as functions of time. Temperatures represented are those observed through a window into the cham¬ 
ber of the gun. Note the large differences between Tr, T b, and T c while the powder is actively burning, resulting from e r 
being greater than unity. (This figure has appeared as Figure 58 in NDRC Report A-323.) 


the pressure, and the emissivity for one round with 
SPDN powder from measurements taken at the 
chamber position. It will be seen that during the 
early stages of burning there is a parallelism between 
the results in the gun and those obtained in the closed 
vessel (Figure 19). The outstanding difference be¬ 
tween the two lies in the course of the emissivity with 
time. Again e r exceeds unity and also exceeds e b , 
hence T R > T B > T c ) however, the emissivity builds 
up faster to saturation in the gun than in the closed 
vessel. This is largely because of the presence of black 
powder in the gun primer; the mass emissivity k for 


black powder gas being much higher than that for 
NH powder gas, the gas in the gun becomes more 
emissive earlier in the shot. 

Figure 21 presents the course of the temperature 
with time, averaged for 13 full-charge rounds, with 
measurements taken at all three positions on the gun. 
It will be noted that the temperatures measured in 
the chamber (position 1) are higher than the other 
two throughout the round, and that, hence, there is a 
large gradient of temperature between the chamber 
and the bore. There is a similar but not so pronounced, 
gradient of pressure. The temperature gradient is 


CONFIDENTIAL 






























POWDER GAS RADIATION AND TEMPERATURE 


51 


greater than can be accounted for quantitatively on 
simple ballistic theory. 65 ’ 113 It is believed that transfer 
of heat to the walls must be largely responsible, inas¬ 
much as heat transfer between the forward portions 
of the gas and the walls takes place much faster than 
that between the forward and rear portions of the gas 
itself. Because of the high opacity of the gas, and the 
high rate of heat transfer to the walls, the temperatures 
measured are those of a relatively thin layer of gas 
cooler than the average, near the window; at the for¬ 
ward holes, this effect is likely to be much more pro¬ 
nounced than in the chamber, and hence the observed 
temperature gradient may be more apparent than 
real. 113 

Firings in a Jet Propulsion Motor 

Firings in an experimental jet propulsion motor, 
23^ in. inside diameter and 9 in. long, equipped with 


two diametrically opposite windows placed near the 
nozzle end were made with six different rocket pow¬ 
ders, both salted and unsalted. The pressures ranged 
up to 3,000 psi. Measurements were taken with a two- 
color photoelectric pyrometer through one window, 
and spectra were photographed through the other, on 
stationary or moving plates, with a small Bausch and 
Lomb quartz spectrograph. 

The experimental findings showed that with salted 
powders even at fairly low pressures of gas, and with 
unsalted powders at pressures above 1,000 psi the 
radiation was near-black-body in distribution, and 
characteristic of temperatures considerably lower 
than the (isobaric) adiabatic temperature. At low 
pressures the spectral features, photographed from 
7000 A to shorter wavelengths, were emission lines 
and bands due to resonance transitions in K, Na, Ca, 
CaO, Fe, Cu, and CuH, together with a continuous 
background. No nonresonance lines or bands were 



TIME TO EJECTION IN MSEC 

Figure 21. Gas temperature versus time curves averaged for 13 rounds at 100 per cent charge with SPDN powder in a 
3-in. gun. The temperatures were observed at three positions: T h chamber; T 2 , 10 in. ahead of origin of rifling; Tz, 51 in. 
ahead of origin of rifling. Note the large apparent gradient of temperature between hole 1 and holes 2 and 3. (This figure 
has appeared as Figure 58 in NDRC Report A-323.) 


CONFIDENTIAL 





























PROPERTIES OF POWDER GAS 


found, nor any due to free radicals, such as CH, C 2 , 
XH, and CN. At all pressures after the early stages 
of firing, strong reversal was observed at the positions 
of Na 5896 A and K 4046 A, as well as in some cases 
at the CaO bands and Ca lines (the plates used were 
not sensitive beyond 7000 A, hence the K 7665- 7699- 
A doublet was not observed). The reversal of these 
intense lines is indicative of a cooled layer of gas near 
the windows; it was strongest when the windows were 
recessed and not flush with the inner wall of the 
motor. 

As already stated, the temperatures observed were 
much below the adiabatic; as an extreme case, un¬ 
salted ballistite 14400, T 0 (isobaric) = 3200 K, gave 
T c about 2700 K at 1,000 psi, windows flush, and 
about 2300 K at 2,100 psi, windows recessed l^ in- 
The emissivity in the red, e r , tended to be greater than 
C6, as in the results for the closed vessels and for the 
3-in. gun. There was also observed considerable fluc^ 
tuation of both the color temperature and the emissiv¬ 
ity even when the pressure was moderately steady. 
This was due presumably to the great turbulence of 
gas in the motor. 

The early course of the temperature is of interest in 
connection with the discussion of Section 2.2; at pres¬ 
sures below 100 psi, the temperatures appeared to be 
below 2000 K, and then rose sharply to near the maxi¬ 
mum temperature at the time the pressures reached 
the range 300 to 700 psi. 


2 5 6 Concluding Remarks 

Anomalous Emissivities 

In the foregoing brief discussion of the radiation 
emitted by powder gas, and the determination of 
powder gas temperature from the intensity of its ra¬ 
diation, it has been indicated that, while the gas 
appears to have high mass emissivity (values rang¬ 
ing up to about 100 cm 2 /g for certain powders) and 
hence should in a thick enough layer radiate as a 
black body, there are certain discrepancies observed, 
particularly when the powder is in an active state of 
burning. The persistent tendency toward excessive 
emissivity in the red region of the spectrum, espe¬ 
cially where the burning goes on at moderate and 
high densities when black body conditions should be 
most nearly fulfilled, leads to the supposition that 
some processes other than thermal ones may be in¬ 
volved in the spectral excitation, even though all the 
observed spectra in the visible and near infrared re¬ 
gions appear to have their origin in thermal excita¬ 
tion alone. 

It seems significant that the anomalous emissivi¬ 
ties become normal when the burning is completed. 
It would seem profitable to extend the studies of the 
radiation as far into the infrared as may be practic¬ 
able, in view of the highly intense bands already ob¬ 
served 160 in that region. While at the low pressures 
and short path lengths used in those studies the con- 



Figure 22. Enlargement of a photograph of the luminosity and turbulence of powder gas in a closed vessel, taken with 
a rotating drum camera on Super XX film. The time scale is referred to the instant of release of pressure by rupture. 
Powder—FNH-M2; loading density—0.18. Image of the slit is superposed. 





CONFIDENTIAL 




POWDER GAS RADIATION AND TEMPERATURE 


53 


tinuous radiation intensity was very low in compari¬ 
son with that of the observed bands, it is possible 
that the bands, if supra thermally excited, may re¬ 
main more intense than the continuous background 
even at gun pressures. 

The errors in temperature introduced by depar¬ 
tures of the emissivities of powder gas from unity are 
evident from equations (48) and (49), which are ob¬ 
tained by recasting equations (42), (43), and (44). 

1 1 _ Xilnei 

~Si T~ c 2 

1 1 _ + X r Xfe In € r /€b 

~T C ~^T~ C2(X r - X 6 ) 

Thus, if e b = 1, e r > 1, the blue brightness temper¬ 
ature is the correct temperature, the red, S r (or T r ) 
is too high, and T c is too low. As an orienting calcula¬ 
tion taking X r = 9000 A, X& = 4500 A, T = 2500 K, 
€r = 2, € 6 = 1, we find T b =T = 2500 K, T r = 2800 K, 
and T c = 2250 K. This would be an extreme case, as 
red emissivities as high-as 2 are not frequently 
found. 


(48) 

(49) 


Absorption of Radiation 

Errors may occur in calibration, and in evaluating 
the records. These are of minor magnitude, however, 
and depend upon the care with which the work is 
done. A more serious possibility of error lies in the 
absorption of radiation by dirt deposited on the win¬ 
dow. A study of this has shown that such deposits 
appear to be re-evaporated in successive firings, and 
condense only after the gas has cooled appreciably. 44 
The deposits do not appear to be spectrally selective, 
and would not, therefore, affect adversely the deter¬ 


mination of the color temperature. It is recommended 
that fresh windows be used in each firing, neverthe¬ 
less, in order to minimize any errors from this source. 


Gas Turbulence 

It must be remembered that powder gas is rela¬ 
tively opaque to radiation coming from considerable 
depths in consequence of the high mass emissivities 
k, as given by equation (41). As a result, the radia¬ 
tion characteristics measured are those of a relatively 
thin layer of gas near the window, which may not be 
representative of the bulk of the powder gas insofar 
as its temperature is concerned because of cooling by 
the walls, and uneven turbulence. 44 The effect of tur¬ 
bulence is strikingly demonstrated in Figure 22. It is 
largely responsible for the individual minor varia¬ 
tions of the observed temperature presented by Fig¬ 
ure 19. It is likewise responsible for the lack of strict 
reproducibility of temperature in guns; thus, when 
the radiation in the 3-in. gun was observed simulta¬ 
neously at two diametrically opposed holes in the 
positions 1, 2, or 3 in a firing, only the major varia¬ 
tions in temperature were found common to both, 
the minor fluctuations differing by as much as 100 K 
at any instant. 65 


Temperature Gradient 

The observational difficulties pointed out here 
may, similarly, make the apparent gradient of tem¬ 
perature between the chamber and bore of a gun 
greater than the real gradient associated with gas 
flow and heat transfer in the gas and between the gas 
and the walls. 


CONFIDENTIAL 





Chapter 3 

INTERIOR BALLISTIC CALCULATIONS 

By William S. Benedict a 


si INTRODUCTION 

T he subject of interiorballistics, interpreted broad¬ 
ly, would include all the phenomena that take 
place inside a gun when it is fired. In the more limited 
sense with which this chapter is concerned, interior 
ballistics calculations are those systematized mathe¬ 
matical methods whereby the motion of the projectile 
in the gun is related to the state of the powder and 
its gaseous products of burning, and to the design of 
the gun. More specifically it is concerned with the 
following variables: the travel of the projectile, L, 
the velocity of the projectile, V, the pressure of the 
powder gas, P, the temperature of the powder gas, T, 
and the time, t. These depend upon the dimensions of 
the gun: its chamber volume, v e , its bore diameter, D , 
its length, X m ; upon the projectile: its weight, M, and 
reaction with the bore; upon the powder: its weight, 
C, chemical composition, and granulation; and upon 
other factors, such as ignition and the temperature of 
both the gun and the powder. 

We may mention some of the practical reasons why 
the entire course of these variables should be known. 
The muzzle velocity is one of the most important 
properties of a gun, and should be accurately repro¬ 
ducible. The pressure-travel curve must be known 
both in order to design a gun of the minimum weight 
that will not burst and also in order to determine 
whether it is safe to use a prospective powder in an 
existing gun. The pressure-time and temperature¬ 
time data must be known in considering gas-operated 
rapid-fire devices, and in making calculations for re¬ 
coilless guns. The temperature plays an important 
role in understanding the thermal and chemical pro¬ 
cesses causing erosion. The pressure and temperature 
at the muzzle affect the blast and flash. 

The ultimate aim of interior ballistic theory must 
be to make predictable in every detail the outcome of 
firing any designated gun with any designated ammu¬ 
nition. In view of the large number of variables in¬ 
volved—and we have not named them all—this aim 
is not likely to be achieved. More modestly, however, 
we may hope to so reduce the problem, with the aid 

a Geophysical Laboratory, Carnegie Institution of Wash¬ 
ington, and National Bureau of Standards. 


of reasonable assumptions and approximations, that 
the art of gun operation and design is based securely 
on scientific fact. 

A historical survey of the development of interior 
ballistic theory falls beyond the scope of this chapter. 
A good summary of the earlier advances is given in 
Cranz’s textbook. 144 505 Methods used by the Army 
and Navy in recent years are described in the Service 
textbooks. 509 515 ’ 520 There have been two main lines of 
approach. In one, the theoretical, the aim is to base 
the largest possible number of relations upon the gen¬ 
eral laws of physics and upon independently deter¬ 
mined properties of the powder and gun. In this way 
a complete prediction might be made without firing a 
shot. In the second, or empirical, approach, the ob¬ 
served pressure-travel-velocity relations for various 
guns firing various charges are correlated so as to give 
a good representation of the overall process, without 
inquiring into the detailed mechanism. Such tables as 
those of Le Due, 509 ’ 515 Roggla, 180 and Bennett 553 result 
from the second approach; these are highly useful in 
predicting the behavior of the conventional types of 
guns and powders, firings of which were used in com¬ 
piling the tables. 18 

In making a fundamental attack on the problems 
of hypervelocity and gun erosion, the first approach 
is obviously to be preferred. The various interrelated 
physical processes are to be given the mathematical 
expression that is in best accord with present-day 
theoretical and experimental knowledge. The result¬ 
ing set of equations is then to be solved, with the aid 
of simplifying assumptions, and approximations, in 
as general a manner as possible. The dependence of 
the variables of interest upon a large number of 
parameters is thus obtained. 

A formalized method of setting up and solving the 
equations is referred to as a system b of interior ballis- 


b The system was developed in a series of formal Division 1 
reports (listed in the bibliography as items 26, 30, 33, 35, 37, 
55, 66), later consolidated and revised into a single report. 69 
During the same period of time improvements 353 ’ 354 ’ 374 made in 
the standard British system of ballistics 431 caused it to have 
some of the same advantages as the Division 1 system. At 
Aberdeen Proving Ground, interest in ballistic systems has 
been directed lately toward the development of functions to be 
used with a differential analyzer. 554 


54 


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THE FUNDAMENTAL EQUATIONS OF INTERIOR BALLISTICS 


55 


tics. Such a system has been developed at the Geo¬ 
physical Laboratory as a part of the Division 1 pro¬ 
gram. For brevity we shall refer to this formulation 
as the Division 1 system. 0 Although work of this na¬ 
ture does not lend itself readily to summarization, we 
attempt in this chapter to state the fundamental 
principles and chief achievements of the Division 1 
system, with emphasis upon the physical content 
rather than the mathematical details. Frequent refer¬ 
ences are made to equations and tabular matter that 
cannot be reproduced here, for which the interested 
reader must consult the original papers, particularly 
the consolidated report. 69 

There is nothing radically novel in the Division 1 
system; its advantage over previous systems consists 
primarily in the large number of factors given explicit 
consideration, and related a priori to the powder 
composition. Naturally the use of a large number 
of parameters leads to rather complicated expressions. 
In its complete form the Division 1 system has been 
aimed toward accuracy and flexibility rather than 
convenience. For certain applications, on the other 
hand, ease of use has been achieved by the excision or 
limitation of variables; and it has been possible to pre¬ 
pare various sets of tables that greatly reduce the 
difficulty of solving many types of problems. 


32 THE FUNDAMENTAL EQUATIONS OF 
INTERIOR BALLISTICS 

3,21 The Basic Processes 

What happens in a gun? The charge, consisting of 
a number of small nearly identical grains of smoke¬ 
less powder, of a definite geometrical shape, is ig¬ 
nited. The solid powder liberates gas at a high tem¬ 
perature, To, and a definite rate, dN/dt, that depends 
on the pressure, P. The pressure in turn depends on 
the quantity of gas, N, its temperature, T, and the 
available volume, v, which is initially the volume of 
the chamber, v c , less the volume of the solid powder 
C/p. The pressure exerted on the base of the projectile 
causes the projectile of mass M/g to move down the 
bore with a velocity V, thus increasing the volume. 
The energy acquired by the moving projectile 
MV 2 /2g, as well as heat lost to the walls of the 
chamber and bore, comes from the powder gas, thus 

c This system has been familiarly referred to as the “Hirsch- 
felder System” after Dr. J. O. Hirschfelder, who was in charge 
of the group that developed it. 


decreasing its temperature. The pressure increase due 
to burning of powder to gas is opposed by the two 
effects just mentioned, and a maximum pressure is 
attained, usually when the volume has about dou¬ 
bled. Thereafter the pressure falls, slowly at first, 
then more rapidly after the powder is completely 
burned. The projectile velocity continues to increase, 
but at a decreasing rate, until it reaches the muzzle. 

There are thus four relationships* among the vari¬ 
ables that are of fundamental importance. These are: 
(1) the equation of motion of the projectile, relating 
its velocity and position to the gas pressure; (2) the 
equation of energy, relating the velocity to the gas 
temperature and quantity; (3) the equation of state, 
relating the gas pressure, temperature, quantity, and 
volume; (4) the equation of burning, relating the gas 
quantity and the pressure. 

Figure 1 attempts to portray schematically the na¬ 
ture and interrelation of these four fundamental 
equations. Of the six variables portrayed, only five 
are independent, the gas volume being linked to the 
projectile position and quantity of gas (or solid pow¬ 
der). We might also consider the time as a primary 
variable in place of the projectile position or velocity. 
By combining the four fundamental equations among 
the five variables, it is possible to arrive at a single 
equation between two variables. It is found most 
convenient to use the velocity as the primary inde¬ 
pendent variable. In the following sections we will 
state the four fundamental equations in their most 
general form, and then indicate how, with the aid of 
certain assumptions and approximations, these are 
cast into a form suitable for combining into the single 
equation, and the nature of the solution of that equa¬ 
tion, in the Division 1 system. 


The Equation of Motion 


The equation of motion is Newton’s law, which 
states that the acceleration of the projectile times its 
mass equals the net force exerted upon it. The accel¬ 
erating force is the difference between the pressure 
exerted by the gas on the base of the projectile, Px, 
and the retarding pressure due to friction, P r , multi¬ 
plied by the area of the bore, A. The acceleration 
of motion may thus be put in the form of equation (1). 


(Px - P r )A = 


MV dV 
g dX' 


(l) 


In the succeeding equations we are less concerned 
with P x than with P, the average pressure of the 


CONFIDENTIAL 




56 


INTERIOR BALLISTIC CALCULATIONS 



Figure 1. The four fundamental equations of interior ballistics are interrelated by the properties of the powder gases and 
by the velocity and position of the projectile. 


powder gas. These are not identical, because a pres¬ 
sure gradient will develop in a column of gas of which 
one end (at the breech) is stationary, and the other 
(at the projectile) is accelerated. In a gun there may 
also be pressure waves due to irregular ignition. A 
complete solution of the problem of pressure and en¬ 
ergy distribution of the powder gas moving in a gun 
according to the laws of hydrodynamics has not been 
found. 69 - 113 The theory due to Kent, 500 which must be 
a good approximation to the truth after the projectile 
is well advanced in the bore, and which has experi¬ 
mental confirmation (Section 6.2.4), relates the pres¬ 
sure drop to the ratio of the weight of charge to that 
of the projectile through equation (2). 


P = 


p (M + C/ 8) 
r x -v?- 


(2) 


The factor 8, depending on C/M , is very close to 3 
except for hypervelocity guns. Although equation (2) 
may not be exact in the early stages of projectile mo¬ 


tion, both because of the approximate nature of the 
Kent theory and the increased importance of ignition 
pressure waves at that period, it is the best available 
relation and is incorporated in the Division 1 sys¬ 
tem. 26 

The other uncertain term in equation (1), the fric¬ 
tion P r , is discussed in more detail in Chapter 6, in 
which equation (1) is a variant of equation (1) of 
Chapter 3. It is there pointed out that P r is in general 
not a simple function of any principal variable. It is, 
however, large compared to P x only at the start of 
motion when the projectile is being engraved; there¬ 
after it is comparatively small. The assumption ex¬ 
pressed by equations (3a) and (3b) is, therefore, 
made in the Division 1 system (and generally in 
other systems). In these equations 

Pr = Px(Px < Po), (3a) 

Pr = ] ~ c -P(Px > Po), (3b) 

c, a constant of the order of 0.04, and P 0 , the starting 


CONFIDENTIAL 





























THE FUNDAMENTAL EQUATIONS OF INTERIOR BALLISTICS 


57 


pressure, are adjustable parameters of the particular 
firing in question. 

With these approximations, and defining m, the 
effective mass of the firing, by 

m = (M + C/5) (1 j (4) 

the equation of motion may be put in the form 


mV dV 
F A dX' 


(5) 


3 - 2,3 The Equation of Energy 

The equation of energy is simply the law of con¬ 
servation of energy, as applied to the interior ballistic 
process, as expressed in equation (6). 

MV 2 

Erei = ~2g —b Eg + Ef + h. (6) 

The individual terms in this equation are discussed 
below. 

1. The first one, E Teh is the energy released when 
the burned powder gas cools from its adiabatic 
“flame temperature” To (defined in Section 2.4.3 as 
the temperature that would be attained if all the en¬ 
ergy of the chemical rearrangement was expended in 
heating the powder gas) to the actual temperature T, 
which is lower than To because energy is expended in 
the processes summarized by the terms on the right- 
hand side of equation (6). The dependence of E re i on 
the chemical composition of the powder, and on the 
density, state of chemical equilibrium, and tempera¬ 
ture of the gas, has been discussed in outline in Sec¬ 
tion 2.4. Calculations of the type described and illus¬ 
trated there show that E Te i is nearly independent of 
density and nearly a linear function of T. In the Divi¬ 
sion 1 system these approximate properties are as¬ 
sumed to hold generally, which leads to equation (7), 

E, re, = Iff G4T = - j), (7) 

in which F, the “impetus” (called “force” in most 
earlier ballistic systems), is defined as nRTo, n being 
the number of moles of gas generated per unit weight 
of powder, and R the gas constant. The ratio of 
specific heats y = C p /C v , is defined as nR/C v . 

2. Of the expended energy terms, MV 2 /2g, the 
kinetic energy of the projectile, is the largest. The 
energy of rotation of the projectile and of recoil of the 
gun are proportional to it, and relatively small, and 
may be included by adjusting M. 


3. E 0 is the kinetic energy of the powder and pow¬ 
der gas. Its exact value is slightly uncertain. A good 
approximation is given by equation (8), which is 
based on the same Kent solution for the motion of 
the powder gas that leads to the assumed pressure 
distribution. 


E a = 


CU 2 

2g8‘ 


(8) 


4. E f is the energy expended in overcoming fric¬ 
tion. Under the assumptions (3) for the friction, it is 
given by equation (9). 

* - f F - a - TT~J pdx 

^ (M +cy ^ 

5. h is the energy lost as heat, principally by forced 
convection from the gas, to the walls of the chamber 
and bore. It is discussed in detail in Section 5.2. The 
theoretical analysis of the problem leads to the result 
that h is approximately proportional to the travel L, 
which in turn is nearly proportional to the kinetic 
energy. Hence the assumption is made that the heat 
loss is proportional to the kinetic energy, 


h = 



( 10 ) 


Defining the effective ratio of specific heats, y, by 
equation (11), 


y - 1 = (y - 1)(1 + 0), (11) 


incorporating the assumptions of equations (7) to 
(10) into equation (6), and remembering the defini¬ 
tion in equation (4), the energy equation becomes 

NF (l-X) = - 1) mV\ (12) 


3 2 4 The Equation of State 

The equation of state of powder gas is discussed in 
Section 2.4.2. Again, it is necessary to recast the 
equation (10), developed there as the most accurate 
equation, into a form more suited to ballistic calcula¬ 
tions. The Abel equation (13) 

p (i~ v ) = nRT = % (13) 

is the same as equation (10) of Chapter 2 except that 
only one of the virial coefficients is retained. In the 
ballistic equation the covolume 77 is considered to be 
independent of density and temperature, and is set 


CONFIDENTIAL 






58 


INTERIOR BALLISTIC CALCULATIONS 


approximately equal to 0.96' in order to compensate 
for the effect of the higher terms in b' at the average 
density prevailing in guns , 69113 0.2 g/cm 3 = 0.0072 
lb/in. 3 . 

The density A depends on the weight of gas N, the 
weight of charge C, the density of solid powder p, the 
chamber volume v c = AXo, and the position of the 
projectile, X = Xo + L, as expressed by 


A = 


N 

AX — (C — N)/ P ’• 


(14) 


Defining the loading density A 0 = C/v c , and the 
term a = (77 — l)/p, the equation of state becomes 


/Z _ Ao _ aA 0 N \ = NFT 

\X 0 p C ) To • 


3 2 5 The Equation of Burning 

The laws governing the burning of propellants are 
discussed in Section 2.2.3. It is assumed that each 
surface of each grain is ignited simultaneously, and 
that the burning of each surface proceeds normal to 
the surface, at a rate dependent on the average pres¬ 
sure. As a consequence, the fraction of powder 
burned, N/C, is related to the fraction/ of the orig¬ 
inal web thickness W remaining unburned. The rela¬ 
tion, known as the form function, depends solely on 
the geometry of the grain. It may in general be ap¬ 
proximated by a quadratic function (16). 

N = k«- hf+ (16) 


In the simplest case (sheet, long strips, or long 
single-perforated grains) of constant-burning-surface 
powder, N/C = 1 — /. In our derivation of the 
equations we shall confine ourselves to this case, 
although the equations for grains of any shape have 
been worked out in the Division 1 system. With 
seven-perforated powders, it is necessary to use two 
sets of constants k, the first until splintering occurs, 
the second to describe the burning of the splinters. 

The experimental evidence adduced in Section 2.2 
led to equations (17a and b) for the pressure-depend¬ 
ence of the burning rate: 


or 


— W ^ = a 4 - bP, (17a) 

(17b) 


The value of the exponent n in equation (17b) is 
from 0.8 to 0.9. Neither equation (17) can be fitted 


into a ballistic system without great computational 
difficulties. Accordingly in the Division 1 system 
equation (17b) is used, with n — 1. The equation of 
burning, for constant-burning-surface powder, thus 
becomes 


dN ^ CBP 
dt W ' 


(18) 


326 The Fundamental Equation 
and Its Solution 

The four fundamental equations, in their simplified 
forms, equations (5), (12), (15), and (18), may now 
be combined. There will be different equations for the 
interval of burning, when N < C ; and after burning 
is complete, when N = C, and equation (18) no 
longer is needed. 


Interval of Burning 

During the interval of burning, we first combine 
equations (5) and (18), and integrate, obtaining 
equation (19). 

N = No + j^-V. (19) 


The constant of integration No is the quantity of 
powder burned at the time of start of projectile mo¬ 
tion ; by the definition (3a) and the equation of state 
(15) it depends on the starting pressure Po according 
to equation ( 20 ). 

No = P °* (1 ~ p^ p ) . (20) 

t -f- aro 


By simple linear substitutions we may now elim¬ 
inate P, T , and N from equations (5), ( 12 ), (15), and 
(19), leaving the fundamental ballistic equation ( 21 ) 
for the interval of burning (for the case of constant 
burning surface). 


1 1 m v 2 = FC— + —— V — — 

2 C ^ AW V AX 0 


A 0 (^- 

\ P 


+ 


aN 0 

C 


amBV 

+ ~AW 


V- 

/ dX 


( 21 ) 


This is a differential equation in which the velocity V 
is the independent variable and the position of the 
projectile X the dependent variable. All the other 
quantities that enter into the equation are constants 
of the firing. Its solution is facilitated by introducing 
the following abbreviations, z, y , q, and r, expressed 
by equations (22) to (25). 


CONFIDENTIAL 







THE FUNDAMENTAL EQUATIONS OF INTERIOR BALLISTICS 


59 


1 . Z, a dimensionless variable proportional to the 
velocity: 

A W 

< 22) 

2 . y, a dimensionless variable related to the travel: 



3 . q, a dimensionless parameter related to the 
starting pressure: 

A Wo ro . 

9 “ FC 2 m{B/WY ' K ‘ ’ 

4. r, a dimensionless parameter related to the gas 
imperfection: 

r = \FCm(B/AW) 2 a. (25) 


With these abbreviations, equation (21) becomes 

dy Z(y - rZ) 

dZ q+Z -$(y -1)/Z*- 

The solution of this linear differential equation is 

y = J( 1 - a) + r{Z - S), (27) 

g +Z-|(f-l)/Z 1 ' (28) 

s - jj'j m 

The integrals J and 8 have been evaluated numer¬ 
ically and tabulated 69 for a large range of the param¬ 
eters q and u = |(y — 1 ), and for closely spaced 
values of the independent variable, the reduced ve¬ 
locity Z. In addition, closed expressions have been 
given by which they may be calculated for other 
values of the parameters. 

The other variables during the interval of burning 
are related in a straightforward manner to the vari¬ 
ables Z and y, the tabulated functions J and S, and 
the parameters r, s, and u. They are: 


where 

J = 

and 


the pressure 




D m C 2 F 2 i 

( B N 

y dz 


1 » 

II 

\W; 

}Z W 


m C 2 F 2 1 
v c A 2 1 

(C 

j 2 q + Z - uZ* 

' /(I - a) - rS 

(31) 

the velocity 

FC 

B „ 


V = 

~A~ 

w z ’ 

(32) 


the time since the start of motion 

- ™ ( V(I L _ n ^L i. [ z m 

vJo P A 2 WJo P ; (36) 
the gas temperature 

T=T '( l -fvz)’ < 34) 

and the fraction burned 

N _ No.rnBV _ No , mFCt V „ _ 
C - C + AW C + A 2 \wj {6b) 


Equations (30) through (35) express all the variables 
in terms of the independent variable Z, which runs 
from zero to its value when all the powder is burned 
(N/C = 1); from equations (19) and (32) this is 


(1 - Np/OA 2 
b mFC(B/W) 2 


(36) 


All functions at the point of complete burning are 
denoted by the subscript b. The maximum pressure, 
usually reached before the end of the burning inter¬ 
val, occurs at the point 

Z P = y(l + yP,). (37) 


All functions at the point of maximum pressure are 
denoted by the subscript p. Although Z p depends on 
the maximum pressure P p , the term ( a/F)P p is small 
compared to unity, so that equations (37) and (31) 
may readily be solved to find the maximum point and 
pressure. If the solution yields Z p greater than Z& it 
means that the maximum point occurs at Z 6 . 


Interval after Burning 


The travel 


L= V f[y + a -l] 

= £[(J - 1)(1 - a) + r(Z - S)]; (30) 


The fundamental ballistic equation (38) for the 
interval after burning is derived by combining equa¬ 
tions (5), (12), and (15): 

^=A»F 2 = FC — v c m(~ o ~ (38) 


CONFIDENTIAL 










60 


INTERIOR BALLISTIC CALCULATIONS 


Its solution, the velocity-travel relation, is 

]■ (39) 

in which Q, the constant of integration, depends on 
the velocity and position at the point of complete 
burning, as expressed by equation (40), 

(40) 


The other variables in this interval are relatively 
simple functions of Q and X; as given by equations 
(41), (42), and (43): 



t = t b + ^- 


__ rx/Xo— tjAo 

V 1(7 2CF - Jx./xiLl. ^ 


Q X i-r)-i/yx, 


T = 



(41) 

(42) 

(43) 


The values at the time the projectile reaches the 
muzzle, denoted by the subscript m, are readily ob¬ 
tained by using the above equations and the overall 
length of the gun, X m . 


3 2,7 Discussion of the Solution 

Nature of the General Solution 

The equations of the preceding section, together 
with the tables for J and S, permit the calculation of 
the complete ballistic solution for any gun using any 
powder of constant burning surface, in terms of the 
parameters F, y, rj, p, Ao, To, Po, m, B, and W. For 
powders of other granulation, the solution is func¬ 
tionally very similar, except that the additional pa¬ 
rameters ko, k i, and A *2 are involved; with seven- 
perforated powders there are two sets of A-’s, the sec¬ 
ond being used in the additional burning period, that 
of the powder splinters, during which the fundamen¬ 
tal velocity parameter varies from Z b to Z s . Because 
of the large number of parameters, and the complica¬ 
tion of the equations, it is rather difficult to evaluate 
the importance of any one variable. The importance 
of each is to some extent brought out in succeeding 
sections; continued use of the system in various bal¬ 
listic applications is, however, needed for full appre¬ 
ciation of the factors. 


In Section 3.3 we present the suggested methods 
whereby numerical values of the parameters may be 
obtained. In Section 3.4 are given methods whereby, 
by limitation of some of the parameters to their most 
usual numerical values, the influence of other param¬ 
eters, which depend upon the choice of the gun di¬ 
mensions and powder type, may be more clearly seen. 
These are illustrated by numerical examples. In Sec¬ 
tion 3.5 are given some applications of these simplified 
methods to problems of design of hypervelocity guns. 


Simplest Possible Particular Case 

We may point out here how the solution in the 
burning interval becomes particularly simple if two 
additional restrictions are made on the parameters: 
(1) If the covolume of the powder gas equals the spe¬ 
cific volume of the solid powder (about 17 in. 3 /lb), 
the constant a and its associated parameter r equal 
zero. This results in the elimination of the integral S 
from the solution; (2) if the starting pressure and its 
associated parameter q is zero, the integral J is much 
simplified, as expressed by equation (44). 

J = (I - uZ)- l ' u . (44) 


Under these conditions all velocity-travel, pressure- 
travel, and temperature-travel curves are alike, and 
their dependence on the gun and powder parameters 
can be more readily seen. In particular, we have 
equation (45) for the pressure-travel curve. 

p _ m C*F*(B\ (1 ~ yy-jf-W-* 

Fe A 2 \wj ( 7 - 1)/2(1 - r/Ao) • V 


The reduced volume at the time of maximum pres¬ 
sure is then dependent only on y, as given by equa¬ 
tions (46) and (47): 


Vv = 


/ 2 y 

\7 + 1/ ’ 


(46) 


, _ m C 2 F 2 (B Y 1 (7 - 1)(t + 1). m) 

p Fc A 2 VlF/l-ryAo 4 t 2 (2t/7 “ 1) 2( " 1} ^ 


Since the functions of y in equations (46) and (47) do 
not vary greatly over the range of values of y usually 
encountered, to this degree of simplification all pres¬ 
sure-travel curves are similar. 

This much-simplified ballistics has been used in a 
number of Aberdeen Proving Ground reports 196 ’ 204 
and was also used by Nordheim 48 in his calculations 
of the heating of guns, as described in Section 5.4.1. 


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EVALUATION OF THE PARAMETERS IN THE BALLISTIC EQUATIONS 


61 


Table 1 . Nominal values of powder constants. 


Powder 

Pyro 

FNH 

(85/10/5) 

NH 

(87/10/3) 

FNH-M2 
(20% N.G.) 

IMR 

(Small arms) 

To (K) 

2320 + 7/W* 

2480 

2425 + 10/W 

3560 

2800 

C v (cal/g-deg.) 

.3653 - .00022/W 

.3456 

.3520 - .00025/W 

.3439 

.3460 

n (moles/g) 

.0466 - .00007/W 

.04482 

.0461 - .0001/W 

.03873 

.04301 

F (ft) 

300,000 + 500/W 

310,000 

309,000 + 500/W 

384,000 

334,000 

7 

1.252 - .0002/W 

1.258 

1.256 

1.223 

1.246 

v (in. 3 /lb) 

29.70 

30.57 

30.49 

26.95 

29.32 

p (lb/in. 3 ) 

.0560 

.0567 

.0571 

.0596 

.0578 

a (in. 3 /lb) 

13.72 

12.93 

12.98 

10.17 

12.02 


* (W = powder web in inches.) 


33 EVALUATION OF THE PARAMETERS 
IN THE BALLISTIC EQUATIONS 

331 Powder Constants 

The powder parameters—the impetus (“force”) F, 
the adiabatic flame temperature T 0 , the ratio of spe¬ 
cific heats t, and the covolume rj —are calculable 
from the chemical composition of the powder. The 
assumptions which make this distinctive feature of 
the Division 1 system possible (in earlier systems 
such constants were estimated from closed-chamber 
measurements) have been outlined in Sections 2.4.3, 
3.2.3, and 3.2.4. In case the exact composition of a 
particular powder lot is not available, or in making 
general calculations, it is useful to have “nominal” 
values of the powder constants, based on the average 
composition of the particular type of powder in ques¬ 
tion. Table 1 presents such recommended values. 
Since the volatile and moisture content of certain 
powders varies with the web, the constant W enters 
into the results. 

For a comparison of the approximate methods of 
calculation of powder parameters with more exact 
methods, and with experimental results, the original 
papers 6 ’ 23 ’ 24 ’ 43 - 113 - 116 must be consulted. Except pos¬ 
sibly for times near ejection, when the gas composi¬ 
tion is uncertain, the nominal parameters will be 
quite accurate. 

3,32 Heat and Friction Losses 

The parameters y, m, and P 0 , which take into ac¬ 
count the energy losses and travel delay due to heat¬ 
ing of the gun and friction between projectile and 
bore, may not always be known a priori. If ballistic 
results on a given gun are available, it may be possi¬ 
ble to adjust these parameters so as to obtain the 
best fit. A closer discussion of the processes of heat 
loss and friction, and of some typical numerical re¬ 


sults, is given in Chapters 5 and 6 . The following 
generalizations from these results may be used to 
estimate the parameters when no experimental knowl¬ 
edge is available. 

The factor /?, the ratio of heat loss (prior to ejec¬ 
tion) to kinetic energy, which relates the effective y 
to the true y [equation ( 11 )] is a function principally 
of the caliber of the gun, being lower for larger guns. 
It also increases with the flame temperature of the 
powder, and with the length of the gun. It, in general, 
decreases as the maximum pressure and muzzle ve¬ 
locity are increased by increasing the density of load¬ 
ing in an existing gun. Some typical theoretical 
values 48 are shown in Table 2 . 


Table 2. Theoretical values of ballistic heat loss, typical 
guns. * 


Gun 

Powder 

0 

7 

Caliber .50 

IMR 

0.40 

1.35 

37-mm 

FNH-M1 

0.27 

1.33 

3-in. 

FNH-M1 

0.25 

1.32 

4.7-in. 

FNH-M1 

0.19 

1.31 

8-in. 

FNH-M1 

0.17 

1.30 

16-in. 

FNH-M1 

0.13 

1.29 


* The values of 8 have been taken from Table XXVII of NDRC Report 
A-262 48 and the values of y have been computed from them by equation 
(11). (Compare Table 2 in Chapter 5, in which the total heat loss to the 
gun tube for these same guns is presented.) 

The most recent experimental evidence (for the 
37-mm and 3-in. guns at Carderock ) 106 would indi¬ 
cate that the values of Table 2, and those given by 
the following formula [equation (48) ] recommended 
in an early report , 30 

. 0.09(7o— Ts) ,/3 CLD~ ilz 

13 = Ws ’ (48) 

are from 10 to 25 per cent too high. For an average 
powder and an average caliber gun (from 37 mm to 
5 in.), 7 will be close to 1.30. This rounded value is 


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62 


INTERIOR BALLISTIC CALCULATIONS 


Table 3. Approximate values of burning rate constant, typical powders. 


Powder 

FNH-M1 

NH 

Pyro 

IMR 

FNH-M2 

B (in. per sec/psi) 

.00033 

.00036 

.00036* 

.00040 

.00050 


* For W < .05 in.; W = .00034. When W > .05 in., W = .00042. 


generally used in the construction of many tables in 
the Division 1 system, and in all following calcula¬ 
tions of this chapter, unless otherwise specified. It 
may be used without introducing gross errors for all 
except the smallest caliber guns (caliber .50 or 
smaller). 

The parameter P 0 cannot be directly correlated 
with its physical meaning of the starting pressure. 
The over-simplification of friction and of the burning 
law destroy the exact correspondence of P 0 with the 
results of static push tests, w hich may roughly be ex¬ 
pressed as a function of caliber in accordance w r ith 
equation (49), 

Po = 2500 D- 1 ' 3 psi, (49) 

which may be used if firing data are lacking. 

It is also permissible to take a starting pressure 
corresponding to No/C = .01. The latter assumption 
is used in most of the following calculations of this 
chapter. When pressure-travel data are available, P 0 
should be determined from them. In general, in addi¬ 
tion to the variation with caliber, P 0 decreases w r ith 
increased loading density and increased run-up of the 
projectile before engraving starts. 

The parameter m, the effective mass, depends upon 



Figure 2. The dependence of 8 on C/M. (This figure has 
appeared as Figure 15-2 in NDRC Report A-397) 


tw r o auxiliary parameters, the friction factor c, and 
the pressure-drop factor 8. Factor c may normally be 
taken equal to .04. It may be somewhat higher in 
high-velocity guns, and also w r hen the velocity is re¬ 
duced in normal guns by reduction of charge; and 
somewhat lower in shorter guns. Factor 8 may be 
taken as 3 in all cases unless the ratio of charge to 
projectile weight exceeds 0.5; for very large ratios 
e = C/M its variation is as shown in Figure 2, based 
on the Kent-Hirschfelder theory. 

3 3 3 Burning Rates 

The burning constant B is of prime importance in 
determining the maximum pressure, as is clear from 
equation (47). It is, however, difficult to correlate 
values of B determined from closed-chamber firings 
with those obtained in guns. Hence whenever pos¬ 
sible B should be determined from the observed maxi¬ 
mum pressure or muzzle velocity. Lacking such data 
the values of B presented in Table 3 may be used. 
These correspond to a P p of the order of 40,000 psi 
and N 0 /C = .01. For higher maximum pressures and 
starting pressures the nominal B decreases somewhat. 
B varies with the initial temperature of the pow^der; 
the tabulated values refer to 25 C. The average tem¬ 
perature coefficient dB/dT is about 0.004 per degree 
centigrade. 

34 SIMPLIFIED NUMERICAL METHODS 
OF BALLISTIC CALCULATIONS 

3 41 Construction of Tables 

The solution of the ballistic equations, as given in 
Section 3.2, followed the happenings within the gun 
in logical sequence from the start of burning to shot 
ejection. In a very large number of ballistic problems, 
however, one final result, the muzzle velocity, and 
one intermediate occurrence, the maximum pressure, 
are the two quantities of greatest interest. These may 
either be given, as the most readily determinable 
measurements, or are the most important results to 
be calculated. It is therefore desirable to simplify the 
ballistic system in such a way that these two quanti- 


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NUMERICAL METHODS OF BALLISTIC CALCULATIONS 


63 


ties may be most readily correlated with the other 
parameters. This entails some recasting of the equa¬ 
tions. Then, by fixing certain of the parameters, tables 
may be constructed which facilitate the solution of all 
the principal equations. 

A complete description of the methods is given in 
Chapters IV and V of Interior Ballistics Consolidated ; 69 
Tables III - XV of that report comprise the principal 
results. They are based on the following values of the 
parameters: Heat loss, y = 1.30 {u = 0.15); Starting 
pressure, N 0 /C = . 01 ; Powder constants, those given 
in Table 1 of the present chapter for nominal FNH- 
M1 powder (the tables may readily be adapted to 
other powders, since only the ratio F/a enters into 
the equations); Granulation, two cases are consid¬ 
ered, ( 1 ) constant burning surface, ( 2 ) seven-perfo¬ 
rated grains, with a ratio of grain length to diameter 
of 2.5, and grain diameter to perforation diameter of 
10 . 

There are five fundamental parameters or auxiliary 
variables that enter into the simplified tables. These 
are: 

1 . The reduced velocity at the completion of burning, 



(for constant-burning grains), (50a) 



(for seven-perforated grains) • (50b) 

2. The pressure function: 

P„ =( - ) • ~P = 2.3969 X 10 ' % P. (51) 

\a/Mi F F 


3. The travel function: 


_ aA 0 
* (X/Xo) - rjAo 


4. The density-of-loading function: 

* = d/Ao (-) 1 Ip) ' 


(52) 

(53) 


5. The velocity function for the period after burn¬ 
ing is complete 


r(Z 6 ,P p °) = (1-.1485Z*,) &-»•*. (54) 

In terms of these functions the velocity after burn¬ 
ing, including the muzzle velocity, is 

F2 = l5m (1 - r?S) ' <»> 

Since q, r, and u have been fixed by assigning definite 


values to P 0 , a, and y, the point of maximum pressure, 
Z p , is a function only of the maximum pressure, as 
expressed by equation (56). 

Zp = f3 (1 + 3 4707 X 10 ^ p »°)- (56) 


Hence the entire ballistic solution is fixed when P p 
and either <f> or Z b is specified. For example, there is a 
relationship among the three principal variables for 
the constant-burning case, stated in equation ( 57 ). 


.99 
ZbJ p 


r e , 2.8763X 10 6 (. 0101 Z b +Z P - .15Z P 2 ) 

L s ' + p? 


] 


(57) 


The simultaneous solution of equations (56) and 
(57) is given in tabular form, one variable being tabu¬ 
lated as function of pairs of arguments: <j>(Z b ,Pp), 
Z b (<t>,P p ). The travel at the end of the burning period 
is given by equation (58). 


\ S b 

z b 


(58) 


From this equation and equation (54) T, and hence 
V m , is related to P p , Z b , and <j>. Additional tables ex¬ 
press r as a function of pairs of arguments. By an 
extension of these methods the pressure, velocity, and 
temperature may be found as functions of time as 
well as of travel of the projectile. 


342 Generalized Pressure-Travel Curves 

Equations (30) and (31) together give the pressure- 
travel curve for the interval of burning, parametri¬ 
cally in terms of Z, and are exact, but involve a num¬ 
ber of constants. If, as in the preceding section, the 
constants No/C, y, and the powder constants 77 and p 
are fixed, the relation between the pressure, in terms 
of its ratio to the maximum pressure P/P p , and the 
travel, in terms of the volume expansion ratio X/X 0 , 
depends only on the loading density A 0 and the fun¬ 
damental burning parameter Z b . The latter moreover 
enters only in the less important terms q and r. Hence, 
if one substitutes in these terms for Z b the value Z in 
the region of the pressure-travel curve from the maxi¬ 
mum pressure to the completion of burning, and the 
value at the maximum, Z v , for lower values of the 
travel, a family of pressure-travel curves are obtained 
which are approximately valid under all conditions, 
and which depend only on one parameter, the load¬ 
ing density. 


CONFIDENTIAL 







64 


INTERIOR BALLISTIC CALCULATIONS 


The most applicable family of such curves, drawn 
for 7 = 1.30, Nq/C = 0.01, and the values of rj and p 
characteristic of FNH-M1 powder (the curves are 
nearly equally valid for other powders) are shown in 
Figure 3. The loading density A 0 is in units of g/cm 3 . 
In order to use these curves the gun constants must 
be known, together with the values of P p and Z b ; if 
the latter are not given with the problem, they may 
be determined from the loading density and muzzle 
velocity by the methods described in the preceding 
section. These curves apply up to the value X b /Xo 
corresponding to Z b , when the burning period is 
complete. 

In the region from X b /X 0 to the muzzle, X m /X 0 , 
the pressure, relative to its value at X b /X 0 , is, accord¬ 
ing to equation (41) dependent only on A 0 , y and 77 
being fixed. The general family of pressure-travel 
curves for this region, for the same values of Ao as 
were used in Figure 3, are shown in Figure 4. They are 
applied by suitable vertical adjustment of the pres¬ 
sure axis to join with the curve from the burning 
interval at X b /X 0 . 

3 4 3 Numerical Examples 

The original reports describing the Division 1 sys¬ 
tem are liberally studded with numerical examples, 
in which each step in the solution of the various types 
of problems considered is worked out in detail. d We 
choose to illustrate the methods by calculations based 
on the 3-in. gun at Carderock, on which extensive 
ballistic measurements were made, as described in 
Chapter 4 and further discussed in Section 6.3. 

The gun constants are: 

A = 7.30 in. 2 ; v c = 223.1 in. 3 ; 

L m = 126.2 in.; M = 12.79 lb. 

The constants for NH powder were: 

C = 4.02 lb; To = 2636 K; F = 321,800 ft; 
rj = 30.06 in. 3 /lb; p = 0.0571 lb/in. 3 . 

The web of the seven-perforated grain is 0.0355 in.; 
we shall carry out the simplified computation in 
which a constant-burning-surface grain is considered, 
with web W = 0.0452 in. 


d In addition to the general reports listed in footnote b, 
separate reports were issued concerning the free-run-up of a 
projectile 29 and recoilless guns. 36 Also some features of the 
Division 1 system were used extensively by investigators in 
Section H of Division 3 in calculations of the ballistics of 
rockets. 163 


The following are the principal derived constants: 
Ao = — = .01802 lb/in. 3 ; 

Vc 

Xo = J- = 30.56 in.; 

^ = 1 + = 5.1296; 

Xo Xo 

1 

a = rj - 

p 


= 12.54 in. 3 /lb; 

, „ 4 M + C/3.14 

m = 1.04--- 

9 

= 0.4548 lb-sec 2 /ft (assuming c = .04). 

First Computation. To calculate the burning con¬ 
stant B from the muzzle velocity V m = 2,700 fps, 
using the assumptions and simplified method outlined 
in Section 3.4.1. From equation (53), 


12.54 

0 (55.50 - 17.52) 

From equation (52), 


0.3302. 


(59) 


fro — 


12.54X.01802 


(5.1296 - 0.5417) 

From equation (55), 

„ 1 - (.15 mVJ/CF) 


0.3 


0.04925. 

= 1.5189. 


(60) 

(61) 


The value of P p ° corresponding to the above values of 
4> and T is located in Table VI of Report A-397; it is 
35,420 psi. For the powder in question this gives, 
according to equation (51), 

^ = lS^9 = 37 - 9201b/inA (62) 


The value of Z b corresponding to the above values 
of P p ° and <f> is located in Table IV of Report A-397; 
it is 1.0397. 

From equation (50a) 


B = W (.99 A 2 /CFm Z b y 2 

= .000420 in. 3 /lb/sec . (63) 


Second Computation. Using the value of B just 
derived, to calculate the complete ballistics, by the 
detailed methods of Section 3.2.6: By assumption, 
u = 0.15 and N 0 /C = 0.01. From equation (23) 

a = .01802 (1/.0571 + .125) = 0.3179. (64) 


From equation (24) 

q = .01 Z b /.99 = 0.0105. (65) 


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NUMERICAL METHODS OF BALLISTIC CALCULATIONS 


65 



Figure 3. Generalized pressure-travel curves, interval of burning. (This figure has appeared as Figure 6-1 in NDRC Re¬ 
port A-397) 



Figure 4. Generalized pressure-travel curves, interval after burning. (This figure has appeared as Figure 6-2 in NDRC 
Report A-397) 


CONFIDENTIAL 











66 


INTERIOR BALLISTIC CALCULATIONS 


From equation (25) 

r = .99a Ao/Zb = 0.2152. (66) 

The remainder of the computation is summarized in 
Table 4. The principal parameter in the burning in¬ 
terval, Z, runs from 0 to Z b = 1.0397. In the after- 
burnt interval the principal parameter, X/X 0 , runs 
from X b /Xo = 2.1404 to X m /X 0 = 5.1296. The aux¬ 
iliary parameter Q is found from equation (40), 

Q = (1 - .1485 Z h ) (2.1404 - 0.5417) 0 - 3 
= 0.9732. (67) 

The results, for the pressure, velocity, temperature, 
and fraction burned, are plotted against the time to 
the muzzle in Figure 5. The pressure and velocity are 
plotted against travel in Figure 6. 

Third Computation. The detailed calculation is re¬ 
peated, using the same values of B and the same para¬ 
metric assumptions, except that the actual seven-per¬ 
forated grain with W = .0355 in. is considered. Inas¬ 
much as the equations for this case have not been 
presented in this chapter, we will give no details. The 
procedure is similar to that shown in Table 4, except 
that the first period ends with the splintering of the 
grains, at Z 8 — 0.8886, N/C = 0.8487, and is fol¬ 
lowed by a second burning period. The results, as 
pressure-travel and velocity-travel curves, are shown 
in Figure 6. 


It is not within the scope of this chapter to discuss 
the correspondence between such theoretical curves 
as those given in Figures 5 and 6 with the experimental 
findings, such as are illustrated in Figures 2 and 3 of 
Chapter 4. The agreement of the pressure-time curves 
is good but far from perfect, because the theoretical 
curves are based on fixed values of the parameters Po, 
c, and t, and an oversimplified burning law. Further 
numerical comparisons of the Division 1 theory with 
ballistic experiments are given in two reports. 113 - 116 

35 INTERIOR BALLISTICS OF 

HYPERVELOCITY GUNS 

351 Introduction 

The equations and tables previously presented and 
discussed in this chapter are quite general. According 
to them, and to common sense considerations as well, 
for a gun of given caliber the muzzle velocity may be 
increased in a number of ways. We may increase the 
length X m ; increase the charge C ; decrease the mass 
of the projectile M ; or increase the impetus of the 
powder F. 

In this section we present some of the contributions 
that the Division 1 system has made to the theory of 
design of hypervelocity guns. Three questions are con¬ 
sidered. First, is there any theoretical limit to the 


Table 4. Illustrative numerical example of complete ballistic calculation, 3-in. gun.* 


Quantity— 

Z 

J 

S 

L 

P 

V 

t 

T 

N/C 

Units— 




in. 

klb/in. 2 

ft/sec 

msec 

K 






Interval (a)— 

Burning 






0 

1.0000 

0.0000 

0.00 

1.01 

0 

0.00 

2636 

0.0100 


.1 

1.0786 

0.1044 

1.62 

10.06 

165 

2.53 

2600 

0.1052 


.2 

1.1864 

0.2198 

3.78 

17.66 

329 

3.31 

2561 

0.2004 


.4 

1.4517 

0.4908 

8.88 

28.74 

658 

4.19 

2482 

0.3909 


.6 

1.7932 

0.8291 

15.14 

35.03 

987 

4.83 

2403 

0.5813 


.8 

2.2342 

1.2567 

22.89 

37.47 

1317 

5.39 

2324 

0.7618 


1.0 

2.8068 

1.8036 

32.63 

37.05 

1646 

5.95 

2245 

0.9622 


1.0397 

2.9399 

1.9298 

34.85 

36.71 

1711 

6.06 

2229 

1.0000 

Source— 


Table 

Table 

Eq (30) 

Eq (31) 

Eq (32) 

Eq (33) 

Eq (34) 

Eq (35) 


X/X 0 


Interval (b)—After-burnt 





Source— 





Eq. (41) 

Eq. (39) 

Eq. (42) 

Eq. (43) 



2.1404 



34.85 

36.71 

1711 

6.060 

2229 



2.5417 



47.11 

27.50 

1993 

6.614 

2084 



3.0417 



62.39 

20.58 

2224 

7.218 

1949 



3.5417 



77.67 

16.23 

2385 

7.771 

1845 



4.0417 



92.95 

13.29 

2508 

8.290 

1762 



4.5417 



108.23 

11.17 

2605 

8.788 

1693 



5.1296 



126.20 

9.35 

2698 

9.352 

1624 



*The value of a used in calculating this table was 0.3134 instead of 0.3179 as given by equation (64). 


CONFIDENTIAL 












INTERIOR BALLISTICS OF HYPERVELOCITY GUNS 


67 



Figure 5. Theoretical ballistic curves, 3-in. gun at Carderock. 


velocity that may be attained in a gun, laying aside 
all thought of practicality? Second, in a gun of given 
total volume firing a projectile of given weight, what 
are the theoretically optimum conditions to produce 
the maximum velocity? Finally, what practical con¬ 
siderations govern the use of modification of the 
theoretical optimum in gun design? 


netic energy of the projectile, MV 2 /2g. As may be 
seen from Figure 2, as C/M approaches infinity, 8 
approaches a finite limiting value, (3y — 1)/(y — 
1 ). 

The equation for the muzzle velocity (39) may be 
rewritten as equation (68), where m and s are defined 
by equations (69) and (70), respectively. 


3 5 2 The Upper Limit to Muzzle Velocity 

The existence of an upper limit to muzzle velocity 
is required, because as the charge-to-mass ratio (C /M) 
increases, the kinetic energy of the powder gas, 
CV 2 /28g, becomes large by comparison with the ki- 


VJ = 


200gF A 0 (l - c)(l - ) 

(7-l)(X m /X 0 )(M - s) 
100M 
M " AX m ' 
lOOAo 

* S(XJX 0 )• 


( 68 ) 

(69) 

(70) 


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68 


INTERIOR BALLISTIC CALCULATIONS 



Figure 6. Theoretical pressure-travel curves, 3-in. gun at Carderock. 


As M approaches zero, the muzzle velocity ap¬ 
proaches the limit given by equation (71). 

vj -» 2(1 ~ c)i ? f- a (i - rjv -1 )- (71) 

(T - 1) 

Using the limiting value for 5, and considering the 
most favorable case of an infinitely long gun (f OT = 0) 
with no friction or heat loss, the limiting muzzle ve¬ 
locity Fii m is expressed by equation (72). 

7-1 

Equation (73), 26 shows the relation between the 
limiting velocity for a projectile and V 8 , the velocity 
of sound in the gas at the adiabatic flame tempera¬ 
ture, which would be the velocity of efflux of a gas 
through an orifice into a vacuum. 



Assuming the F corresponding to FNH-M2 powder 
in equation (72) leads to a value of 36,400 fps for 
the limiting muzzle velocity. Friction and heat losses 
increase with increasing velocity to such an extent 
that the actual limiting muzzle velocity for even the 
hypothetical projectile of zero mass probably would 
not exceed 20,000 fps. For an actual gun firing a 
projectile of useful weight, of course, the possible ve¬ 
locity is much less. 498 (See Section 33.2.2.) 

3.5.3 The Optimum Conditions 

Considering equation (68) from a more practical 
angle, it will be observed that the muzzle velocity 
depends on n, X m /X 0 , P p , A 0 , y, and the powder com¬ 
position. The main problem in gun design is the 
determination of the best values of these quantities 
that will yield a desired muzzle velocity. It is clear 
that for any values of the other quantities, /x can be 


CONFIDENTIAL 























INTERIOR BALLISTICS OF HYPERVELOCITY GUNS 


69 


chosen to give any velocity up to the impractically 
high limit. Tables 69 have been prepared, based on 
the simplified ballistic tables described in Section 
3.4.1, that shows the values of n required to give ve¬ 
locities up to 5,000 fps at various values of P p , 
X m /X 0 , and A 0 , for a typical single-base and a typical 
double-base powder of constant burning surface. An 
illustrative portion of the tables, for V m = 4,000 fps, 
is reproduced as Table 5. The variation of n with V m 
is presented in Table 6. Figures 7 and 8 illustrate the 
dependence of the velocity on the loading density and 
on the volume ratio for various fixed values of m- 

The most efficient gun at a given velocity and cali¬ 
ber has the highest value of /z (greatest kinetic energy 
of projectile, least length of gun). Conversely, at a 
given n it is desirable to maximize the muzzle velocity. 

The optimum density of loading, (A 0 )i, is therefore 


defined as the density of loading that gives the high¬ 
est muzzle velocity for a given powder, at a pre¬ 
assigned value of the maximum pressure, projectile 
mass, and total gun volume. It is obtained by differ¬ 
entiating equation (68) with respect to A 0 and equat¬ 
ing to zero. Since T, f m , and $ are all functions of A 0 , 
the resulting equation is a complicated function of 
P p , X m /X 0 , and n/ (/z + s) = M/(M + C/8 ); it may 
be solved numerically. (A 0 )i has been computed and 
tabulated under the usual simplified ballistic assump¬ 
tions. In a precisely similar manner, one defines and 
obtains the optimum volume ratio (X m /Z 0 ) 2 . 

The double optimum, (A 0 ) 3 and (X m /X 0 ) 3 , is defined 
as the combination of loading density and gun length 
that gives the maximum muzzle velocity for any 
given value of the maximum pressure, projectile mass, 
total gun volume, and powder type. It is obtained by 


Table 5. Values of /i = 


lb/in. 3 for V m = 4,000 fps. 

A 2i m 


Powder 

FNH-M1 

FNH-M2 

A 0 , (g/cm 3 ) 

0.6 

0.7 

0.8 

0.6 

0.7 

0.8 

X m /X 0 



P p = 50,000 psi 



3 

.6026 

.6386 

.5982* 

.6744 

.6641 

.6126* 

4 

.6091 

.6746 

.6814 

.7046 

.7462 

.6917* 

5 

.5716 

.6431 

6726 

.6714 

.7383 

.7235 

6 

.5281 

.5989 

.6376 

.6260 

.6904 

.7041 

8 

.4512 

.5158 

.5589 

.5406 

.6047 

.6348 

10 

.3917 

.4495 

.4916 

.4722 

.5322 

.5669 




P p = 60,000 psi 



3 

.6645 

.7456 

.7670 

.7662 

.8202 

.7909* 

4 

.6508 

.7464 

.8049 

.7666 

.8523 

.8799 

5 

.6025 

.6961 

.7635 

.7173 

.8087 

.8585 

6 

.5523 

.6404 

.7085 

.6620 

.7318 

.8096 

8 

.4677 

.5440 

.6071 

.5652 

.6466 

.7066 

10 

.4039 

.4705 

.5274 

.4905 

.5633 

.6203 


* Powder not all burned when projectile leaves muzzle. 



Table 6. 

Values of u — 

lb/in. 3 for P p = 50,000 lb/in. 3 and X m /X 0 

= 5. 


V alUuo U1 fl — A V 

AX m 

Powder 

FNH-M1 

FNH-M2 

A 0 , (g/cm 3 ) 

0.6 

0.7 

0.8 

0.6 

0.7 

0.8 

Vm (fps) 







1,500 

4.8594 

5.5005 

5.8426 

5.5688 

6.1204 

6.2046 

2,000 

2.6766 

3.0277 

3.2106 

3.0756 

3.3764 

3.4143 

2,500 

1.6662 

1.8831 

1.9924 

1.9215 

2.1063 

2.1227 

3,000 

1.1174 

1.2613 

1.3306 

1.2947 

1.4163 

1.4211 

3,500 

0.7864 

0.8865 

0.9316 

0.9167 

1.0003 

0.9981 

4,000 

0.5716 

0.6431 

0.6726 

0.6714 

0.7383 

0.7235 

4,500 

0.4244 

0.4763 

0.4951 

0.5032 

0.5452 

0.5353 

5,000 

0.3190 

0.3570 

0.3681 

0.3829 

0.4128 

0.4006 


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70 


INTERIOR BALLISTIC CALCULATIONS 



Figure 7. Curves of velocity versus density of loading 
at constant n for X m /X 0 = 5 and P p = 50,000 psi. (This 
figure has appeared as Figure 7-5 in NDRC Report 
A-397.) 


Table 7. Conditions for the double optimum. 


Powder 

FNH-M1 

FNH-M2 

P P 

(Ao) 3 


(Ao) 3 


(psi) 

(g/cm 3 ) 

(X m /X 0 ) 3 

g/cm 3 

(X m /X 0 ) 3 

10,000 

.284 

4.70 

.248 

4.69 

20,000 

.465 

4.27 

.413 

4.33 

30,000 

.594 

3.93 

.540 

4.05 

40,000 

.689 

3.66 

.640 

3.83 

50,000 

.764 

3.45 

.720 

3.65 

60,000 

.823 

3.29 

.786 

3.49 


simultaneous solution of the maximizing equations 
for (A 0 )i and (X m /X 0 ) 2 . The double-optimum density 
is independent of the muzzle velocity, but (X w /X 0 ) 3 
is weakly dependent on V m . Some values of the double 
optimum are listed in Table 7; the values of X m /X 0 
are the limiting values at V m = 0. The values of the 
mass/volume parameter \x that correspond to these 
double-optimum conditions, in the hypervelocity 
range 3,000 to 5,000 fps are listed in Table 8. 

It must be emphasized that the calculations lead¬ 
ing to these theoretically optimum conditions are 
based upon the simplified ballistic assumptions of 
Section 3.4. They serve therefore only as an orienta¬ 
tion to the approximate dimensions to choose in seek- 



Figure 8. Curves of velocity versus X m /X p at constant 
n for A 0 =0.7 g per cu cm and P p =50,000 psi. (This fig¬ 
ure has appeared as Figure 7-4 in NDRC Report A-397.) 


ing the best design. They must be modified by prac¬ 
tical considerations long familiar to the gun designer, 
as is outlined in the next section. After deciding upon 
the approximate optimum dimensions, calculations 
by the detailed ballistics may be made, and further 
tested for the optimum conditions, by the methods 
sketched here. 


3,5 4 Practical Considerations in the 
Design of Hypervelocity Guns 

Limiting Density of Loading 

There are various considerations that limit and 
modify the attainment in actual practice of the opti¬ 
mum conditions. It is clear from the tables of the 
preceding section that guns with higher maximum 
pressures, which lead to higher muzzle energies, are 
most efficiently operated at high loading densities. 
The loading density cannot, however, be increased 
indefinitely; it is limited by the density of solid pow¬ 
der and by the geometry of the grains and the cham¬ 
ber, to values near 1.0 g/cm 3 . 

At such high densities it is, moreover, impossible 
to operate a gun with any degree of safety, because 


Table 8. Values of n — - lb/in. 3 at the double optimum. 


Powder 

FNH-M1 

FNH-M2 

P P (psi) 

40,000 

50,000 

60,000 

40,000 

50,000 

60,000 

V m (fps) 







3,000 

1.142 

1.382 

1.665 

1.208 

1.477 

1.735 

3,500 

0.788 

0.960 

1.158 

0.841 

1.037 

1.218 

4,000 

0.561 

0.688 

0.832 

0.609 

0.749 

0.884 

4,500 

0.409 

0.499 

0.604 

0.451 

0.552 

0.651 

5,000 

0.302 

0.369 

0.448 

0.344 

0.414 

0.490 


CONFIDENTIAL 



























INTERIOR BALLISTICS OF HYPERVELOCITY GUNS 


71 



Figure 9. Pressure-travel and velocity-travel curves for various shapes of powder grains. (This figure has appeared as Fig¬ 
ure 7-7 in NDRC Report A-397.) 


when the free volume in the chamber is small, the 
pressure and velocity become very sensitive to small 
variations in the web, burning rate, or charge weight 
of the powder. Thus the changes in these properties 
that are inevitable in the processing of powder and 
its storage under field conditions may lead to unsafe 
variations in the pressure. Furthermore, at high load¬ 
ing densities pressure waves and other erratic results, 
usually attributable to ignition difficulties, are fre¬ 
quently encountered. 

Hence the upper practical limit, using present-day 
powders and igniters, can rarely exceed 0.8 g/cm 3 (ex¬ 
cept for small arms). This limit may be increased 
somewhat by improved methods of ignition, such as 
are provided by the use of long primers. 


Limiting Pressure 

As mentioned above, an increase in the maximum 
pressure, limited by the safe strength of the gun, 
which in turn is limited by the properties of the gun 


metal and the tactical considerations that do not per¬ 
mit any superfluous bulk in weapons, is in general 
desirable. Equally important in attaining a high value 
of n is an increase in the effective pressure, which is 
defined as the pressure on the base of the projectile 
averaged over the travel: 


, _ mVrr? .06 VJ 

eft gAL 9 (1 - Xo/XJ^ 


(74) 


The ratio P e s/P p is a convenient measure of the 
efficiency of a gun. At a given density of loading the 
maximum P^/P p occurs at the P p that corresponds 
to the optimum density of loading. Higher pressures 
can be reached only if the pressure-travel curve has a 
peak near the origin of travel, and hence a compara¬ 
tively low P e s/P P - It is therefore inefficient to use a 
maximum pressure much above that for which the 
loading density is optimum. Since at present this is 
limited as just mentioned to about 0.8 g/cm 3 , it is us¬ 
ually not desirable to aim for maximum pressures much 


CONFIDENTIAL 








72 


INTERIOR RALLISTIC CALCULATIONS 


LENGTH OF GUN IN CALIBERS 



Figure 10. Relations between chamber volume and gun length, for a 90-mm gun at P p — 60,000 psi, V m — 4,500 fps; FNH- 
M1 powder. (This figure has appeared as Figure 12-3 in NDRC Report A-397.) 


above 60,000 psi. It is for this reason that Tables 
7 and 8 are not carried beyond that limit. Thus new 
advances in the strength of gun steels, while very 
desirable in reducing the weight of the weapons, 
might not permit a great extension of the present-day 
limitations on ballistic design, unless accompanied by 
an increase in the density of loading. 

Preferred Volume Ratio 

The optimum volume ratio (X w /X 0 ) 3 usually cor¬ 
responds to a comparatively large chamber and short 
travel. However, as is clear from the flat maximum in 
Figure 8, an increase in X m /X 0 by one or two units 
above the optimum value does not decrease the veloc¬ 
ity greatly, and implies a considerably smaller cham¬ 
ber, for the same total volume. This in turn means a 
decreased charge and a more efficient utilization of 
the powder. 

As a typical numerical example, we note from 


Tables 5, 6, and 7 that at V m = 4,000 fps and P p 
= 50,000 psi the optimum loading density for Ml 
powder is 0.764, corresponding to n = 0.688 and 
Xm/Xo = 3.45. At that density an increase to X m /X 0 
= 5.0 would lower n only to 0.661. That is, the same 
muzzle velocity and maximum pressure could be ob¬ 
tained with a decrease in chamber volume and pow¬ 
der charge of 27 per cent, at the expense of an in¬ 
creased total gun volume of only 4 per cent. Thus, 
unless tactical considerations demand either an un¬ 
usually short gun or an unusually light charge, the 
best volume ratio is one or two units above the theo¬ 
retical optimum. 

There is another strong reason for using a gun 
length greater than the optimum in most cases; 
namely, the larger the gun, the lower are the tempera¬ 
tures and pressures of the powder gas when the pro¬ 
jectile leaves the muzzle. Too high muzzle tempera¬ 
tures and pressures result in considerable muzzle flash, 
muzzle blast, and increased velocity dispersion. 


CONFIDENTIAL 




























INTERIOR BALLISTICS OF HYPER VELOCITY GUNS 


73 


LENGTH OF GUN IN CALIBERS 



Figure 11. Relations between chamber volume and gun length at optimum density of loading, for a 90-mm gun at V m = 
4,500 fps; FNH-M1 powder. (This figure has appeared as Figure 12-6 in NDRC Report A-397.) 


Choice of Powder 

The comparison between FNH-M1 and FNH-M2 
powders in the tables in this section shows that the 
double-base powder, with an impetus (F = nRT 0 ) 
greater by almost 25 per cent, gives an increase in 
muzzle energy, under optimum conditions at the 
same maximum pressure and muzzle velocity, of 
about 10 per cent. Because an increase in impetus 
nearly always connotes an increase in gas tempera¬ 
ture, which in turn leads to much greater erosion, it is 
rarely desirable to make use of the slight gain in bal¬ 
listic efficiency provided by double-base powders.® It 
would, however, be clearly advantageous to develop 
powders with a high n but low T 0 . 


e A study at Aberdeen Proving Ground of the dependence of 
muzzle velocity on the potential and on the force of powder 
showed that at optimum conditions the muzzle velocity de¬ 
pends much less on the energy properties of the powder than 
on the maximum pressure or the length of the gun. 204 


Granulation of Powder 

All the previous discussion in this section has been 
based on the simplest powder granulation, that of 
constant burning surface. The maximum ballistic 
efficiency would be obtained from a very progressive 
powder; that is one whose surface would increase with 
burning in such a way that the amount burned would 
increase proportionately with the available volume, 
resulting in a uniform pressure until the powder was 
completely burned, with a correspondingly high 
Peff/Pp- 

A mixed charge, consisting of a small amount of a 
very fine granulation, to burn rapidly and establish 
the constant high pressure, and the remainder of 
single-perforated tubes of very small inside diameter, 
fireproofed to burn only on the inside, would approach 
this ideal granulation. The seven-perforated powders 
in common use are only very slightly progressive; it 
has been shown that they are practically equivalent 


CONFIDENTIAL 

























74 


INTERIOR BALLISTIC CALCULATIONS 



Figure 12. Relations between gun length and muzzle velocity at double-optimum conditions for a 90-mm gun, firing a pro¬ 
jectile of normal weight (24 lb) at various muzzle velocities with either FNH-M1 or FNH-M2 powder. (This figure has 
appeared as Figure 12-11 in NDRC Report A-397.) 


ballistically to constant-surface powder of 27 per cent 
greater web, the advantage of increasing surface 
before splintering being counterbalanced by the rap¬ 
idly decreasing surface after splintering. Cord gran¬ 
ulation, used extensively in British guns, is very de¬ 
gressive, and hence ballistically less efficient; it can, 
however, be loaded and ignited controllably at higher 
densities of loading than other powders, and thus has 
many advantages. 

The effect of granulation on gun performance has 
been calculated. A typical example of the differences 
in pressure-travel and velocity-travel curves resulting 
from changes in grain shape is shown in Figure 9. 

3 5 5 The Design of a 

Hypervelocity 90-mm Gun 

To illustrate the types of calculations described in 
Section 3.5.3, a number of curves have been drawn 69 
for the 90-mm gun described in Section 33.1.4 (A 
= 10 in. 2 ), firing a projectile of 24 lb. We then have 


2400 

M v e XJX o’ 

L = 0.1v c (XJXo - 1). 

Taking the cross-sectional area of the chamber to 
be three times that of the bore, the total length of gun, 
L g = L —f 0.033 v c . Hence from the tables of /jl we may 
plot v c against X m /X 0 or L, at various muzzle veloc¬ 
ities, maximum pressures, and densities of loading. 
Such a curve for V m = 4,500 fps, P p = 60,000 
psi, the listed densities of loading, and FNH-M1 
powder, is reproduced as Figure 10. The optimum 
density of loading is that corresponding to the shortest 
gun at each value of X m /X 0 . The optimum densities, 
for various maximum pressures, are shown in Figure 
11. The shortest gun for each maximum pressure and 
the corresponding density of loading are the double¬ 
optimum conditions. The velocity and gun length 
corresponding to these are plotted, for two powder 
types, and several maximum pressures, in Figure 12. 
It must be remembered that in practice one would use 


CONFIDENTIAL 




















INTERIOR BALLISTICS OF HYPERVELOCITY GUNS 


75 


a somewhat smaller chamber, smaller density of load¬ 
ing, and longer gun than are given by Figure 12; but 
the curves give a good idea of the effect of chang¬ 
ing maximum pressure, powder type, or required muz¬ 
zle velocity, in designing a 90-mm gun. The applica¬ 
tion of these calculations to a specific case is described 
in Section 31.8. 

3 5 6 Interior Ballistics of 

Gun Firing Subcaliber Projectiles 

The preceding discussion, the principal results of 
which are depicted in Figure 12, illustrate how dif¬ 
ficult it is, even with all improvements in theoretical 
and practical gun design, to increase the velocity of 
conventional guns firing conventional projectiles into 
the range of 5,000 fps, unless very long weapons 
are used. 

In order to attain velocities this high without 
lengthening the gun to 100 calibers or more, one must 
find other means of decreasing n. To use a projectile 


that is light for its caliber makes for poor exterior- 
ballistic stability, as is brought out in Chapter 8. If, 
however, a small-caliber projectile can be fired from 
a large-caliber gun, good exterior ballistics may be 
obtained together with a lowered value of n- As the 
tables indicate, if n can be reduced to one-quarter of 
its normal value, by firing a projectile of half the 
caliber of the gun, the velocity increase will be of the 
order of 2,000 fps. 

If the decreased caliber of the projectile is obtained 
by means of a sabot (Chapter 29), there is some de¬ 
crease in ballistic efficiency, in that a portion of the 
energy of the powder is used in providing kinetic 
energy to the sabot that is subsequently discarded. An 
alternate means of obtaining a subcaliber projectile 
with good ballistic form is to fire a deformable pro¬ 
jectile from a tapered-bore gun, as is described in 
Chapter 30. The methods and tables of the earlier 
sections of this chapter may be applied to the guns 
firing these unconventional projectiles. 


CONFIDENTIAL 




Chapter 4 

INSTRUMENTATION FOR EXPERIMENTAL BALLISTIC FIRINGS 

By H. B. Brooks a 


4.1 the role of experimental 
BALLISTICS 

S ome of the problems in interior ballistics that 
face the designer of a hypervelocity gun have been 
outlined in Chapter 3. Their solution requires more 
exact knowledge of the behavior of guns and projec¬ 
tiles during firing than has been available in the past, 
especially under conditions of hypervelocity. Divi¬ 
sion 1, NDRC, working through the National Bureau 
of Standards and the Geophysical Laboratory of the 
Carnegie Institution of Washington, attempted to ob¬ 
tain some of that knowledge by setting up a ballistic 
range and making physical measurements with labor¬ 
atory precision during the firing of guns. 

The present chapter deals primarily with the in¬ 
strumentation developed for this purpose. It comprises 
a general description of the firing range, an outline of 
the ballistic events recorded during the firing of a gun, 
and a condensed description of the apparatus used to 
record them. b A sketch is also given of a separate de¬ 
velopment at the Geophysical Laboratory of tech¬ 
niques and instruments for measuring ballistic quan¬ 
tities directly by means of apparatus attached to the 
projectile while it is moving in the bore of a gun. 

Although certain instruments have been used for 
many years at proving grounds for some routine bal¬ 
listic measurements, many others had to be developed 
especially for these firings. This development was 
carried out during the first two series of firings of a 
3-in. naval gun. Even though this is not a hypervelo¬ 
city gun, a third series of firings was made with it not 
only for the sake of checking the performance of the 
instruments but also for the sake of obtaining bal¬ 
listic data that could be used in elucidating some as¬ 
pects of ballistic theory. The principal results are pre¬ 
sented under appropriate subject headings in Sections 
2.5.5, 5.2.2, 6.3, and 7.4. 

The first hypervelocity gun to which this instrumen¬ 
tation was applied was the 37-mm, T47 (Section 


a National Bureau of Standards and Geophysical Labora¬ 
tory, Carnegie Institution of Washington. 

b Complete descriptions are given in a series of NDRC re¬ 
ports listed as the following items in the Bibliography: 39,44, 
65, 108, 131, 132. 


31.7), firing pre-engraved projectiles at a muzzle ve¬ 
locity of approximately 3,500 fps. These firings repre¬ 
sented the first recorded systematic study of the in¬ 
terior ballistics of a medium-caliber gun with pre-en¬ 
graved projectiles. The results are summarized in 
Sections 5.2.2, 6.4, and 7.4. 

An important feature of the later stage of Division 
l’s program was the development of a hypervelocity 
90-mm gun, firing pre-engraved projectiles at a muz¬ 
zle velocity of 4,000 fps, as outlined in Section 31.8. 
It was planned that eventually one of these guns 
would be used in a series of ballistic firings. As a prep¬ 
aration for that investigation, and especially in order 
that there would be available a firm basis of compari¬ 
son, plans were made for conducting a series of firings 
with a 90-mm gun, M1A1. The gun was procured 
through the cooperation of the Army Ordnance De¬ 
partment and its tube was modified to accommodate 
measuring instruments just as the 3-in. gun had been. 
When Division l’s sponsorship of this work had to be 
withdrawn near the close of 1945, the Bureau of Ord¬ 
nance of the Navy Department supported its con¬ 
tinuance by the National Bureau of Standards. 

4-2 THE CARDEROCK RANGE 

AND THE GUNS 

421 Range and Laboratory 

The ballistic range, including the associated labora¬ 
tory, was constructed for Division 1 on the grounds 
of the Navy Department’s David W. Taylor Model 
Basin at Carderock, Maryland, a few miles from the 
District of Columbia. The range is 500 ft long and is 
horizontal. Concrete piers for the support of veloc¬ 
ity solenoids are located at 50-ft intervals along the 
range. A set of ducts extends along the range from the 
laboratory with a manhole opposite each pier and one 
near the gun. Through these ducts wires extend to the 
laboratory from each pier and from the gun. At the 
distant end of the range there is a sand-filled butt 
with heavy concrete walls and a removable concrete 
block for a cover. 

The laboratory is a one-story frame structure lo¬ 
cated about 50 ft behind the gun. It has two rooms, 


76 


CONFIDENTIAL 



THE CARDEROCK RANGE AND THE GUNS 


77 


one 18 by 18 ft, the other 12 by 18 ft. It is possible to 
make the laboratory light-tight, as is necessary when 
the very sensitive films are to be placed on or removed 
from the recorder drums. A ventilating system op¬ 
erates through the attic, and heat is supplied by a 
system of thermostatically-controlled electric heaters. 
The wiring connections to the laboratory apparatus 
are suspended from hooks in the ceiling. An interior 
view of the larger laboratory room is shown in Figure 1. 

4 2 2 The 3-in. Gun and its Mount 

A 3-in./50-cal. gun, Mark 3, Mod. 7 and mount 
were provided by the Bureau of Ordnance, Navy De¬ 
partment. Very few shots had been fired from the gun 
since it was relined. The primary support for the gun 
was a heavily reinforced concrete slab 20 ft sq and 4 ft 
thick. Steel foundation bolts, cast into the concrete 
slab, extended up through a cushioning layer of 2-in. 
plank and through a piece of 2-in. armor plate to 


which the gun mount was bolted. About 1 ft beneath 
the gun and forward from the mount was a concrete 
pier with a top about 2 by 3 ft. This pier supported 
the apparatus which was connected directly to the 
gun for the purpose of measuring displacement, veloc¬ 
ity, and acceleration of recoil. 

Some modifications were made in the gun in the 
shops of the Taylor Model Basin before the ballistic 
firings were started. The firing mechanism was modi¬ 
fied so that the gun could be fired electrically. Two 
holes were bored into the powder chamber at opposite 
ends of a diameter to allow the pressure and tempera¬ 
ture of the powder gas to be measured at each instant. 
Two bands with projecting arms were fitted to the 
gun so that recoil apparatus could be readily attached 
to the gun barrel. A muzzle plate with a suitable 
clamp was fitted to the gun to permit the instant of 
ejection of the projectile to be determined. 

After eleven rounds had been fired, three additional 
pairs of holes were drilled through the barrel between 



Figure 1. Interior of larger laboratory room at ballistic range. 


CONFIDENTIAL 





78 


EXPERIMENTAL BALLISTIC FIRINGS 


the powder chamber and the muzzle. The two holes 
composing each pair were at opposite ends of a 
horizontal diameter. 

4 2 3 The 37-mm Gun 

One of the experimental 37-mm, T47 guns describ¬ 
ed in Section 31.7 was provided by Division 1 through 
the Franklin Institute. This gun fired a 1.62-lb pre¬ 
engraved projectile at a muzzle velocity of about 
3,500 fps. Its chromium-plated tube had been made 
by boring a 40-mm tube for 37-mm and chambering 
it for a standard 40-mm cartridge case necked down 
to accommodate a 37-mm projectile. 

In order to prepare this tube for ballistic firings, a 
collar about 6 in. long was shrunk on the barrel, near 
the base of the powder chamber, and two holes were 
drilled through this collar into the powder chamber 
to receive powder-pressure gauges. The flash hider 
was replaced with a muzzle plate carrying muzzle 
fingers, an optical ejection indicator, and the antennas 
for the microwave interferometer. 

Six types of projectiles, two types of powder, and 
three types of primers were used in the ballistic 
firings. 

4 3 EVENTS RECORDED AND QUANTITIES 
DETERMINED 

431 Method of Measurement 

The important events which occur when a gun is 
fired are recorded on photographic films as oscillo¬ 
graph traces and in other ways by means of the in¬ 
struments described in Section 4.4. From these rec¬ 
ords the quantities involved in interior ballistic cal¬ 
culations are determined. Some of the events involve 
only time, such as the time the firing pin strikes the 
primer or the time the projectile passes the muzzle of 
the gun. Other events require the determination of a 
quantity as a function of time, such as the temperature 
or the pressure of the powder gases. The measure¬ 
ment of each event requires special consideration. 

4 3 2 Firing Pin Strikes the Primer 

The firing pin of the 3-in. gun carried two spring 
contact devices, normally closed. The free end of each 
spring was loaded to increase its inertia. These con¬ 
tact devices faced in opposite directions. When the 
firing pin started to move, the inertia of one of the 


loaded springs broke its contact and thus recorded 
this event on the oscillograph trace. The inertia of the 
other loaded spring, however, held its contact more 
tightly closed until the firing pin struck the primer, 
when the inertia of its loaded spring opened this con¬ 
tact and thus made its record. The average time of 
travel of the firing pin was about 15 msec, which was 
about 5 msec longer than the time from the ignition 
of the primer to the ejection of the projectile, when 
full service charge was used. 

4 3,3 Start of Powder Pressure 

The time at which the powder pressure started was 
obtained from a film by determining the time at 
which the trace for the pressure in the powder cham¬ 
ber showed the first sign of curving. Because of the 
slow rise of pressure at the beginning, the time at 
which it started could not be determined precisely. 
The uncertainty in the time of start of pressure in the 
3-in. gun did not often exceed 0.2 or 0.3 msec in a 
total time to ejection of about 8 msec. 

4 3,4 Start of Radiation 

The time at which radiation from the powder gas 
started was obtained from its trace on the film by 
determining the time at which the trace started to 
curve. The initial rise in radiation was quite rapid so 
that the time at which it started could be determined 
within 0.1 msec. The start of radiation usually occur¬ 
red slightly before the start of pressure. 

43 5 Start of Recoil 

The start of recoil of the gun was indicated by the 
opening of a contact on the recoilmeter (Section 
4.4.7), which was designed to operate when the gun 
had moved less than 0.001 in. The operation of this 
contact produced a step on the film traces the time of 
which could be determined to 0.01 msec. The start of 
recoil usually occurred before the measured start of 
pressure or of temperature of the powder gas. It is of 
interest to note that the recoil was initiated in the 
3-in. gun by the forward motion of the powder charge 
before the projectile started to move. 

4 3 6 Start of Projectile 

The time of the start of the projectile was original¬ 
ly recorded on the oscillograph film by an inertia con- 


CONFIDENTIAL 



EVENTS RECORDED AND QUANTITIES DETERMINED 


79 


tact-breaking device located within the projectile. 
This device transmitted its starting signal when the 
projectile had moved only a few thousandths of an 
inch. It required a wire stretched between the pro¬ 
jectile and the muzzle. It had to be discarded when 
the microwave interferometer (Section 4.4.15) was 
installed. However, the approximate time at which 
the projectile started to move could be obtained from 
the microwave interferometer method by observing 
when the trace started to curve. 

4 3 7 Travel of Projectile in Gun 

For the 3-in. gun the travel of the projectile was 
observed by three independent methods. In the first 
method the instant was recorded at which the rotat¬ 
ing band made contact with each of several insulated 
pins inserted in radial holes along the gun barrel. This 
method is probably the most precise of the three, but 
is limited by the necessarily small number of holes. 
The second method uses wire-resistance strain gauges 
(Section 4.4.10) to determine the instant at which the 
projectile passes given transverse planes, as discussed 
in Section 7.4.1. Variations in the ratio of band pres¬ 
sure to powder pressure introduce uncertainties in the 
interpretation of results. The third method locates 
the projectile by the microwave interferometer (Sec¬ 
tion 4.4.14). This method gives a large number of 
points along the gun barrel but at the disadvantage 
of requiring elaborate apparatus. It was the only 
method used in determining the motion of the pro¬ 
jectile in the 37-mm gun. 

4 3 8 Velocity of Projectile in Gun 

Several methods have been used experimentally for 
measuring this quantity. The preferred method con¬ 
sists in plotting the curve of projectile position as a 
function of time, as obtained from the data yielded 
by the microwave interferometer (Section 4.4.14), 
and graphically differentiating this curve to obtain 
the curve of velocity as a function of time. 

4 39 Acceleration of Projectile in Gun 

The acceleration-time curve for the projectile has 
been obtained by graphical differentiation of the ve¬ 
locity-time curve. No direct method of measuring the 
projectile acceleration was used in the Carderock fir¬ 
ings. A method tried in a 20-mm gun is described in 
Section 4.5. 


4 310 Muzzle Velocity 

Approximate values of the muzzle velocity for the 
3-in. gun have been determined by two independent 
methods, one of which involves the extrapolation of 
the range velocity backward to the muzzle, the other 
the extrapolation of the velocity in the gun forward 
to the muzzle. The extrapolation required in the first 
method can be more accurately performed, but the 
value obtained for the muzzle velocity requires a cor¬ 
rection for the effect of the blast. This correction has 
not been determined directly, and the uncertainty in 
its value as found by approximate procedures con¬ 
stitutes the chief limitation on the accuracy of the 
method. Presumably the blast accelerates the pro¬ 
jectile, and hence the first approximate value of muz¬ 
zle velocity given by extrapolation of the range veloc¬ 
ity will be too large. The muzzle velocity for the 
37-mm gun was obtained by this method of extra¬ 
polation. 

4 311 Velocity of Projectile on Range 

For both guns the velocity of the magnetized pro¬ 
jectile was measured by recording on an oscillograph 
film the times at which the projectile passed through 
solenoids located at measured positions along the 
range (Section 4.4.11). It was estimated that the 
average velocity between any pair of solenoids 
was determined with an accuracy of at least 1 
fps. 

The observed values for the 3-in. gun extend from 
1,730 fps with 50 per cent of service charge to 2,730 
fps with 100 per cent charge. For the 37-mm projec¬ 
tiles the velocities range from 3,260 to 3,870 fps for 
projectiles weighing 1.92 to 1.34 lb. 

4 312 Deceleration of Projectile on Range 

The deceleration of 3-in. projectiles on the range 
has been computed for each successive trio of sole¬ 
noids, and for the first and the last pair of solenoids. 
Observed values range from 178 ft/sec 2 for 50 per 
cent of service charge to 380 ft/sec 2 for 100 per cent 
of service charge. It was found that for 3-in. pro¬ 
jectiles having identical rotating bands the decelera¬ 
tion a , in ft/sec 2 , is given with reasonable accuracy by 

a = 0.1947—157. (1) 

For the 37-mm projectiles the observed values of de¬ 
celeration range from 600 to 1,250 ft/sec 2 . 


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EXPERIMENTAL BALLISTIC FIRINGS 


4.3.13 Displacement of Gun in Recoil 

The displacement of the 3-in. gun in recoil as a 
function of time has in most instances been directly 
determined by the step-by-step recoilmeter; in a few 
rounds the continuous recoilmeter has been used (see 
Section 4.4.7). Other methods, used in a few cases, 
are (1) the integration of the velocity-time curves of 
the velocimeters, and (2) the double integration of 
the acceleration-time curves of the accelerometers. 
Theoretically, all the records of the motion of the gun 
in recoil can be compared through differentiation or 
integration of the curves drawn by the various in¬ 
struments. 

4 314 Velocity of Gun in Recoil 

The velocity of the 3-in. gun in recoil as a function 
of time has been measured by three different methods, 
namely, (1) direct measurement with a velocimeter 
(Section 4.4.8); (2) differentiation of a displacement¬ 
time curve; and (3) integration of an acceleration¬ 
time curve. The method that has been most extensive¬ 
ly used is the graphical differentiation of the displace¬ 
ment-time curve given by the step-by-step recoil¬ 
meter (Section 4.4.7). However, this method does not 
work well when the velocity and the displacement are 
small, and other methods are at least of equal accura¬ 
cy for the first part of the curve. 

4,315 Acceleration of 3-in. Gun in Recoil 

The acceleration of this gun in recoil has been re¬ 
corded directly by the crystal accelerometer, and by 
differentiators operating on the current from a veloci¬ 
meter (Section 4.4.9). It hasbeenobtained indirectly by 
the differentiation of a velocity-time curve and by the 
double differentiation of a displacement-time curve. 

The acceleration of the gun at a given point con¬ 
sists of two parts, namely, its basic acceleration, and 
the acceleration resulting from vibrations in the gun 
structure. The accelerometers may have a very rapid 
response, indicating all the sudden changes in acceler¬ 
ation (including those resulting from vibrations); the 
graphical-differentiation methods, on the other hand, 
are suitable only for giving average values covering 
about 0.1 msec. Both types of method have their uses. 

Maximum values of acceleration in recoil for the 
3-in. gun range from nearly 7,000 ft/sec 2 for 100 per 
cent of service charge down to about 2,000 ft/sec 2 for 
50 per cent charge. For 100 per cent and 90 per cent 


of service charge the maximum acceleration in recoil 
occurs about 4.5 msec before ejection. For the lower 
values of charge the maximum occurs at longer inter¬ 
vals before ejection, the interval for 50 per cent 
charge being about 7 msec. 

4.3.16 Pressure of Powder Gas 

The pressure of the powder gas was usually meas¬ 
ured at four positions along the 3-in. gun by the 
gauges described in Section 4.4.12. The data from the 
pressure-time curves of each round have been com¬ 
bined with the data from the displacement-time curve 
of the projectile for that round to give the pressure in 
each hole as a function of the position of the projectile. 
The displacement-time curve of the projectile, as 
determined by the microwave interferometer, has 
been used in all rounds when available, but in other 
rounds it was obtained either from strain-gauge data 
or computed from measurements of the recoil of the 
gun. The pressure on the base of the projectile as a 
function of projectile displacement has been deter¬ 
mined from the pressures at the several holes and 
plotted for each round. Figure 2 shows the pressure¬ 
time curve for 100 per cent charge in round 65, as 
measured at each of four holes. Figure 3 relates to the 
same round but shows the pressure as a function of 
projectile travel. 

The values of maximum pressure obtained in the 
various rounds may have been influenced by at least 
three factors, namely, condition of the interior surface 
of the gun barrel, the initial position of the projectile, 
and the temperature of the gun barrel and the powder. 
There is a definite indication that the maximum pres¬ 
sure is lowest with the projectile in its normal posi¬ 
tion and the interior surface of the barrel greased. 
The maximum pressure is somewhat greater if the 
gun barrel is dr}q it is a little greater if the gun barrel 
is greased and the projectile is advanced so that it is 
seated against the rifling; and it is greatest when the 
projectile is advanced and the inner surface of the 
gun barrel is dry. 

Maximum values of powder pressure in the powder 
chamber of the 3-in. gun for service charge ranged 
from 38 to 49 kilopounds psi. Maximum pressures in 
the 37-mm gun ranged from 52 to 64 kilopounds psi. 

4317 Temperature of Powder Gas 

Measurements of the temperature of the powder 
gas were made by observing the intensity of the emit- 


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PRESSURE IN PSI * 10 


EVENTS RECORDED AND QUANTITIES DETERMINED 


81 



Figure 2. Pressure-time curves for 100 per cent charge, 3-in. gun. 



DISPLACEMENT FROM MUZZLE IN FT 

Figure 3. Curves of pressure versus projectile displacement for 100 per cent charge, 3-in. gun. 


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82 


EXPERIMENTAL BALLISTIC FIRINGS 


ted radiation as described in Section 2.5.4. One-color 
and two-color pyrometers (Section 4.4.13) have been 
used on the 3-in. gun in various combinations at hole 
positions 1,2, and 3, as far as the limited number of 
recording channels permitted. Unfortunately, it ap¬ 
pears that the temperature is far from uniform 
throughout the powder chamber or along the bore of 
the gun and the results for any one round at any in¬ 
stant give the temperature of only a small local vol¬ 
ume in front of the window, rather than the average 
temperature at that instant throughout the chamber 
or across the diameter at the position of the hole. For 
this reason, together with the fact that the limited 
supply of recording apparatus restricts the types of 
measurements made in any one round, the typical 
rounds chosen for the report must not be thought of 
as in any sense typical of the complete body of the 
data, but merely as examples. The results of the tem¬ 
perature measurements are summarized in Section 
2.5.5. 

In the holes opening into the powder chamber it 
has been found that for service charge (1) radiation is 
first observed at or shortly before the start of pressure; 
(2) a brief flash of high-temperature radiation (about 
2300 K) coinciding with the first sharp rise of pres¬ 
sure to about 1,000 psi, is followed by rapid cooling 
to about 1900 K, and then by a gradual rise toward 
a maximum; (3) the maximum temperature of 
2550 ± 100 K is reached at about 5 msec before ejec¬ 
tion; the maximum pressure occurs about 1 msec 
after the maximum temperature; (4) as further ex¬ 
pansion of the gas continues, the temperature drops 
rapidly until at ejection it is near the limit of measure¬ 
ment (1600 K). 

4 3,18 Strain Measurements on the Gun 

Strain measurements 65 made by means of the 
gauges described in Section 4.4.10, which were placed 
on the external surface of the 3-in. gun barrel, have 
been of two kinds: (1) qualitative measurements 
which served to determine the time of passage of the 
projectile, and (2) quantitative measurements, used 
in studying the forces acting on the gun tube, in order 
to determine their influence on the friction between 
the gun and the projectile. This evaluation is a diffi¬ 
cult problem, as is brought out in Chapter 7 (especial¬ 
ly Section 7.4). 

4,319 Jump of Gun 

The jump of the 3-in. gun as a function of time was 


measured for all but one of the first 11 rounds. 39 The 
measurements showed that the muzzle of the gun 
started to move upward at about 1 msec before ejec¬ 
tion and had moved nearly 2 mm at ejection. 

4 3 20 Pressure in Recoil Cylinders 

Measurements of pressure in the recoil cylinder of 
the 3-in. gun were made under two conditions. In the 
first, the recoil cylinders were filled with fluid to the 
vent hole, whereupon the pressure began to rise sever¬ 
al milliseconds before ejection. In the later rounds the 
cylinders were first filled and then a small quantity 
M pt in most cases) was removed. Under these con¬ 
ditions the initial rise of recoil-cylinder pressure oc¬ 
curred after ejection. This delay simplified the cor¬ 
relation of the motions of the gun and the projectile 
because the opposing force exerted by the recoil 
cylinders is zero throughout the interval considered. 
The distance the gun recoiled was not changed ap¬ 
preciably by the removal of the liquid. 

4 3 21 Ejection of Projectile 

Two methods of recording the ejection of the pro¬ 
jectile were tried. One depended upon the mechanical 
contact in an electric circuit while the other was an 
optical method. An ejection signal was recorded when 
the front of the rotating band made contact with a 
muzzle finger mounted on the muzzle plate. Complete 
ejection occurred slightly later, as about 2 in. of the 
projectile was still in the muzzle of the gun when the 
signal was recorded. 

An optical system was developed with the expecta¬ 
tion that a signal would be obtained when the nose of 
the projectile intercepted a beam of light passed 
in front of the muzzle and which activated a 
photoelectric cell. As explained in Section 4.4.15, 
this optical method of determining ejection did not 
work well because the beam was obscured by com¬ 
pressed gases and smoke 0.6 to 1.0 msec before the 
nose of the projectile arrived at the beam. 

4 3 22 Pictures of Muzzle Smoke and Flash 

The smoke and gases emerging from the muzzle 
of the gun and the flash were photographed by the 
high-speed cameras described in Section 4.4.16. 
Gases emerging from the muzzle usually obscured 
the projectile so that it did not become visible un¬ 
til it was several feet in front of the muzzle. The pic- 


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INSTRUMENTATION 


83 


tures showed the development of the flash over the 
ball of emerging gases. It usually started at several 
different points at about the same time and then 
merged into a single ball of flame. The different 
powders produced about the same type of flash 
when service charges were used; but the flash devel¬ 
oped with different speeds when reduced charges 
were used. 

4 3 23 Intensity of Muzzle Flash 

The intensity of the muzzle flash from the 37-mm 
gun was measured with the flashmeter described in 
Section 4.4.17. Measurements were made at several 
positions in front of the muzzle. The results show no 
significant correlation with powder classes except for 
the marked difference between M-l powder (low 
flash-intensity) and M-5 powder (high flash-inten¬ 
sity). 

4 3 24 Pictures of Projectile in Flight 

On several rounds pictures of the 37-mm projectiles 
in flight were made about 75 ft in front of the gun by 
illuminating the projectile with a “microflash” (Sec¬ 
tion 4.4.18). In some cases two cameras were placed 
so that the angle between their lines of sight was 90°, 
thus making it possible to measure the yaw of the 
projectile. However, the cameras that were used 
were so small that the measurements of yaw were 
not precise. In other cases the two cameras were 
placed a short distance apart on a line parallel to 
the path of the projectile and stereoscopic pictures 
were obtained. 

4 3 25 Shock Waves 

A shock wave is initiated as a pulse at the muzzle 
of the gun when the projectile is ejected and is propa¬ 
gated outward from the muzzle in the form of a spher¬ 
ical shell. A zone of the sphere lying along that great 
circle which is perpendicular to the sun’s rays refracts 
the rays to produce a shadow of that great circle on 
the ground or other background. Some of these shad¬ 
ows have been photographed by the projectile camera. 
They were not observed in the first examination of the 
pictures made with the 3-in. gun. They were first 
observed in one of the pictures taken in a continua¬ 
tion of this research. A re-examination of the pictures 
from the 3-in. gun showed the shadows of the waves 
quite distinctly. 


44 INSTRUMENTATION c 

441 General Outline 

The instruments located on or attached to the gun 
produce electric signals that are transmitted to the 
laboratory and control the deflection of the electron 
beams of cathode-ray oscillographs. Because a rather 
high voltage on the deflecting plates is required to 
deflect these beams, an amplifier is built into each 
oscillograph to increase the voltage of the trans¬ 
mitted signals. Other electronic equipment between 
the instruments and the oscillographs is mainly for 
mixing circuits or for matching impedance. In all but 
the temperature measurement, the records are made 
on a rapidly moving film mounted on a drum. Tem¬ 
perature measurements are recorded by small cam¬ 
eras rigidly mounted in front of the oscillograph 
screen, with a sweep circuit to produce the horizontal 
motion of the oscillograph spot. 

The timing of the moving film or of the motion of 
the sweep circuit is accomplished by giving to the 
electron beam a small vertical deflection every milli¬ 
second. The speed of the moving film is such that 
these millisecond pulses are about 1.5 cm apart. On 
the film of the still cameras the distance between 
timing pulses is about 9 mm. A common-time signal 
is recorded on each trace when the firing pin begins to 
move (Section 4.3.2). This signal makes it possible to 
correlate the times of events recorded on different 
films. For some of the events only the time of occur¬ 
rence needs to be recorded. Other phenomena require 
the measurement of quantities that vary with time, 
for which purpose two types of recording instruments 
have been used. One gives a continuous record, the 
other a step-by-step record. 

4,4,2 Instruments for Making the Records 

The instruments for making the records were 
oscillographs, recorders, and auxiliary equipment. 
The latter includes the wiring for operating the record¬ 
ers, the devices for making a definite base line on the 
film, and the calibrating apparatus. Figure 4 shows 
an assembly of recording apparatus including two 
details of special parts. 

The two oscillographs are of the DuMont Type 


c The experience of the Geophysical Laboratory in recording 
explosion pressures 3 by means of oscillographs 7 was available 
in planning the installation at Carderock. 


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84 


EXPERIMENTAL BALLISTIC FIRINGS 





DETAILS OF SHUTTER RELEASE MECHANISM 


Figure 4. Recording apparatus for ballistic measurements. 


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INSTRUMENTATION 


85 


208. Directly in front of them is the recorder, which 
is essentially a light-tight housing enclosing an alumi¬ 
num drum 19 in. in diameter with a 4-in. face. The 
drum is belt-driven by a d-c motor. The housing can 
be opened to permit the mounting of a photographic 
film on the drum. An electrically operated shutter 
controls the exposure of the film. The shutter opening 
is wide enough so that the two oscillographs may re¬ 
cord on a single film. A release magnet opens the 
shutter circuit and allows the shutter to close after 
any desired number of revolutions not exceeding 5. 
The optical system requires a lens to focus the spot of 
each oscillograph on the film. 

The drums were synchronized with each other and 
with the firing mechanism of the gun so that the 
shutters of all the drums opened immediately after 
the slot holding the ends of the film passed the shutter. 
One drum was used as a master drum and a contact 
on it operated the firing mechanism. This same con¬ 
tact flashed lights on the other drums which allowed 
the operators to synchronize their drums with the 
master drum by observing the stroboscopic motion of 
a disk mounted on the drum shaft. By this arrange¬ 
ment the record always appeared in the central por¬ 
tion of the film. 

Two type&of records were obtained, one in which 
the time of an event is indicated by an abrupt dis¬ 


placement of the trace, the other, in which the ordi¬ 
nate of the curve at any time is proportional to the 
magnitude of the quantity measured. In Figure 5 
the lowest trace of (A) is an example of the former 
type of record, and the trace of (B) illustrates the 
latter type. For both traces the number above the 
trace is the number of milliseconds from a com¬ 
mon time signal not shown in the figure. The figures 
below trace (A) indicate the steps on a recoilmeter. 
The other traces of (A) will be referred to later. 

4 4 3 Oscillograph Calibrators 

The oscillograph calibrators break up the measur¬ 
ing trace into a definite series of short sections, stair¬ 
step fashion, the ordinate of each step corresponding 
to a definite value of the quantity that has just been 
recorded by the trace. The unique feature of these 
calibrators is the switching mechanism, in which an 
electrically released falling weight opens a series of 
switches, connected to a special network, to produce 
the steps on the film by changes of either resistance 
or voltage. This feature is shown in Figure 6. An im¬ 
portant feature of these calibrators is the fact that all 
switches are opened to produce changes in resistance 
or voltage. This feature of operation insures that no 
chattering of contacts can occur, even when they are 



B 


Figure 5. Types of traces obtained in ballistic measurements with 3-in. gun. 

CONFIDENTIAL 






86 


EXPERIMENTAL BALLISTIC FIRINGS 



Figure 6. Switching device of the oscillograph calibra¬ 
tor. The switches are actuated by a falling weight. 


opened at the rate of one every millisecond. Figure 7 
is a diagram of the circuit of the calibrator used with 
the velocimeter. 

4 4 4 Timing of Film 

The timing of each film is accomplished by record¬ 
ing on it a continuous series of pulses that occur at 
intervals of 1 msec. These timing pulses may be seen 



R ie total resistance in battery circuit when all switches are closed. 
0, 1, 2, ... 5 are switches opened by falling weight, 
ri, rz, . . . rs, are resistances added to circuit by opening of switches 
1,. . . 5 

Eb is voltage of battery and E a that of standard cell. 

G is galvanometer. 


Figure 7. Oscillograph calibrator circuit for veloci¬ 
meter. 


on the traces shown in Figure 5. The same series is 
recorded on all the traces of the films that are exposed 
at any one time. The series on the different traces of a 
set are correlated by having one or more distinctive 
events, called the common time, recorded on each 
trace. The timing pulses for the films are initiated by 
a 1,000-c fork and are impressed on the oscillograph 
circuit by electronic devices called “pulsers.” In the 
course of the investigation these devices have under¬ 
gone improvements which make the pulses appear 
simultaneously on all the oscillographs and increase 
the precision of reading the position of the pulses on 
the film. The description in the next section is limited 
to the improved form as used in the third series of 
firings with the 3-in. gun. 132 

4 4 5 Pulsers 

The timing-pulse generator is designed to convert 
the 1,000-c output of the tuning fork (and related 
assembly) into a series of pulses in which the voltage 
rises sharply, remains nearly constant for about 25 
Msec, then decreases just as sharply to zero. The dis¬ 
tance on the film between the rising side of the pulse 
and the descending side is such that the cross hair of 
the reading microscope may be accurately set mid¬ 
way between the two sides. This condition enables 
accurate settings to be made. An identifying signal, 
inserted in the trace of each film by the firing-pin 
contact, provides a common-time origin for all the 
films. This feature, in connection with the simulta¬ 
neity of the timing pulses, makes possible very 
accurate coordination of the times of all recorded 
events. 


4 4 6 Developing of Films 

The exposed films required special handling and 
developing because of the stringent requirements for 
the finished negatives. 39 The development and initial 
fixing were carried out in the dark because the film 
was too sensitive to permit the use of even red light. 
The observed differential shrinkage ranged from 0.4 
per cent to nearly 1 per cent and the resulting 
error in time measurements did not exceed 5 Msec 
and was usually less. 

4 4 7 Recoilmeters 

Two forms of recoilmeters have been used to meas¬ 
ure the displacement of the gun in recoil (Section 


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INSTRUMENTATION 


87 



TO 

OSCILLOGRAPH 


Figure 8. Recoilmeter for measuring amount of recoil of a gun. 


4.3.13). Both operate on the same general principle, 
namely, the variation of resistance in a circuit by the 
motion of a sliding contact. In the first recoilmeter 
the motion of the gun drives the slider along over a 
series of silver contact segments between which resist¬ 
ors are connected. Figure 8 shows this instrument 
with a part cut away to show details. The step-by-step 
record obtained with this instrument gives a precise 
indication of the time at which the displacement of 
the gun attains certain values. The slider of this re¬ 
coilmeter is connected to the gun by a steel tube. Pro¬ 
vision is made for indicating the start of recoil (Sec¬ 
tion 4.3.5) by the separation of two contacts, one 
attached to the slider and the other to the base of the 
recoilmeter. 

The second recoilmeter operates on the same gen¬ 
eral principle but the series of discrete contacts is re¬ 
placed by a slide wire. Because of the form of the rec¬ 


ord given by this instrument, it is called the contin¬ 
uous recoilmeter. Its slider is carried by the dovetail 
slide of the step-by-step recoilmeter, and the slide- 
wire assembly is fastened to the side of the base of the 
step-by-step recoilmeter. 

4 48 Velocimeters 

Two forms of velocimeters have been used for 
determining the velocity of recoil of the gun (Section 
4.3.14). In the first one used in this investigation the 
linear motion of the gun causes the armature of a 
small a-c generator to rotate. This requires inter¬ 
mediate mechanical means to convert the linear re¬ 
coil of the gun into rotary motion of the armature. In 
the second form the linear motion of a coil of wire in a 
radial magnetic field is nominally identical with the 
motion of the gun and the mechanical connection 


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88 


EXPERIMENTAL BALLISTIC FIRINGS 


of the velocimeter to the gun is therefore much 
simpler. 

Rotary Velocimeter 

The rotary velocimeter is a magneto-type bipolar 
a-c generator constructed to have a substantially 
linear relation between angular velocity and electro¬ 
motive force over about 110° of each half-cycle. Figure 
9 is a view of the generator. The stationary field mag¬ 
net is composed of five Alnico magnets of the type 
used in General Electric Type DP-9 d-c ammeters. 
The armature, which is a single narrow coil wound on 
a bakelite tube, revolves around a stationary iron core 
located coaxially between the polepieces of the mag¬ 
nets and supported by pedestals attached to the ends 
of the base plate. The ends of the armature coil are 
connected to silver slip rings mounted on the bakelite 
tube. With no current in the armature, no armature 
reaction can occur, and because the armature core 
does not rotate, there is no reaction from it that can 
affect the field flux either in magnitude or distribution. 
The instantaneous induced electromotive force is ac¬ 
curately proportional to the instantaneous angular 


velocity, and the wave form for any constant angular 
velocity is independent of the magnitude of this 
velocity. 

Because the generator must be mounted and used 
near the gun, and consequently be exposed to outdoor 
temperatures, the dependence of its induced electro¬ 
motive force on the temperature of the generator is 
important. It was found that an increase of 1 centi¬ 
grade degree in its temperature lowered its induced 
electromotive force by 0.014 per cent. 

The armature of the generator is rotated by the 
gun by means of two thin steel tapes that wind in 
grooves in a sheave attached to the tubular armature 
“shaft.” As the gun recoils, one tape unwinds from 
the sheave and the other winds up. During counter¬ 
recoil the functions of the two tapes are interchanged. 
If vibrations in the driving system of the armature are 
neglected, the angular velocity of the armature of the 
generator is proportional to the linear velocity of the 
gun. The velocity of the gun was obtained from the 
mechanical dimensions of the driving system and the 
electrical calibration of the generator. The electrical 
calibration was obtained by rotating the armature at 
a known uniform speed and determining the gener- 



CONFIDENTIAE 





























INSTRUMENTATION 


89 


ALNICO MAGNET 



GUN BRACKET 


CORE SLEEVE 


TOP HALF OF CROSS-SECTION 
SHOWING DIRECTION OF 
MAGNETIC FLUX 



MAGNET —1Z7/EZ?- 


Figure 10. Linear velocimeter for measuring velocity of recoil of a gun. 


ated electromotive force during the portion of the 
cycle when it was uniform. Thus the relationship be¬ 
tween the recorded electromotive force and the veloc¬ 
ity of the gun was obtained. 

Linear Velocimeter 

In the linear velocimeter, shown in Figure 10, a 
helical coil of wire, which can move only in its axial 
direction, is located in the radial magnetic field be¬ 
tween an iron rod and the inner surface of an outer 
coaxial iron tube. The coil is wound in an annular 
groove cut in a bakelite tube that slides on the iron 
rod. The magnetic flux is set up by an Alnico magnet. 
This flux is uniform in the axial direction for about 
6 cm. This uniformity was checked by observing that 
equal deflections of a ballistic galvanometer were 
obtained when the coil was suddenly moved through 
small equal distances throughout this region. These 
deflections along with known constants of the system 
were used to determine the voltage which was induced 
when the coil was moved with unit velocity. The 
magnetic structure of the accelerometer is fastened 
to the gun pier and the bakelite tube is attached to 
the gun barrel in such a manner that when the gun 
recoils, the velocimeter coil is drawn through the 
magnetic field without any physical contact between 
the coil tube and the iron structure. 


4 4 9 Accelerometers 

Differentiators for Velocimeter 

In addition to the use of the crystal accelerometer 
described later in this section, gun acceleration was 
measured with instruments which are in effect devices 
for differentiating the electromotive force generated 
by the velocimeter. Two types have been considered 
and tried, one of which employs a mutual inductance, 
the other a capacitance in series with a resistance. 

Figure 11 is the circuit diagram of the mutual- 
inductance accelerometer. The principle involved is 
that the voltage induced in the secondary of a pure 
mutual inductance is directly proportional to the 
time-rate of change of current in the primary. Except 
for some disturbing elements, the output voltage of 
the generator is directly proportional to the velocity 
of recoil. The output voltage of the generator is fed 
into the grid circuit of a 6J7 tube, the plate circuit of 
which includes the primary of the mutual inductor. 
The tube is so biased that, for the range of input 
signal voltage used, the waveform of the plate current 
is a faithful reproduction of the waveform of the 
input voltage. Hence the current in the primary of 
the mutual inductor has the same waveform as the 
voltage induced in the velocimeter. The voltage in¬ 
duced in the secondary of the mutual inductor is di- 


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90 


EXPERIMENTAL BALLISTIC FIRINGS 


GENERATOR OF VELOCIMETER 



Figure-11. Mutual-inductance accelerometer for meas¬ 
uring acceleration of recoil of a gun. 


rectly proportional to the time-rate of change of the 
plate current, and hence the deflection of the oscillo¬ 
graph at any instant is theoretically proportional to 
the acceleration of the gun. However, difficulties 
were encountered in the use of this method during 
the firings of the first 11 rounds. 39 

These difficulties were caused by distributed capac¬ 
itances in the air-core mutual inductor used. Since 
then, this inductor has been replaced by an inductor 
with a magnetic core which had a winding of fewer 
turns and hence smaller distributed capacitance. A 
filter was connected between the inductor and the 
oscillograph to suppress high-frequency oscillations. 
The resulting improved differentiator has been used 
as an accelerometer to obtain a continuous record of 
acceleration from a voltage that at every instant is 
proportional to the velocity of recoil. It has also been 
used to record the velocity of recoil from the drop in 
potential across a slide-wire, namely, the continuous 
recoilmeter described in Section 4.4.7. Satisfactory 
records have been obtained. 

The type of differentiator using a capacitance in 
series with a resistance was used on a very limited 
number of rounds so that insufficient data were ob¬ 
tained to determine its usefulness. However the 
mutual inductance type had the advantage of being 
much more simple to construct and use than the ca¬ 
pacitance-resistance type. For this reason the mutual 
inductance type was used in most cases. 

Crystal Accelerometer 

The crystal accelerometer consists of a piezoelectric 
crystal having its electrodes connected to a capacitor. 


The crystal is so cut and mounted that the quantity 
of electricity developed on its electrodes is propor¬ 
tional to the acceleration to which the crystal and 
any associated load are subjected. The potential drop 
over the parallel capacitor is equal to the quotient 
obtained by dividing the quantity of electricity de¬ 
veloped by the sum of the capacitances of the crystal, 
the line, and the parallel capacitor respectively. The 
accelerometer impresses on the oscillograph a voltage 
which is partly the result of the instantaneous acceler¬ 
ation at the point of attachment to the gun and part¬ 
ly the result of the vibrations of the crystal and those 
of the gun, each at its individual resonant frequency. 
As the voltages caused by the acceleration of these 
resonant vibrations may at any instant be several 
times as large as the voltage produced by the non- 
vibratory acceleration of the gun, it is necessary to 
use a filter in the oscillograph circuit to suppress the 
voltages produced by the resonant vibrations. 

The crystal accelerometer embodied a crystal unit 
made of a newly developed material,* 1 ammonium 
diphenyl phosphate (ADP). The details of its con¬ 
struction and attachment to the gun, the method of 
calibration of the crystal, and a block diagram of con¬ 
nections of the crystal, cathode follower, filter and 
oscillograph have been described. 65 

4 410 Strain Gauges for Projectile 
Displacement in the 3-in. Gun 

As the projectile passes any given section of the 
gun, the barrel expands. If the time of occurrence of 
this expansion is recorded for a number of points 
along the barrel, a displacement-time curve can be 
plotted. A record of barrel expansion, as obtained in 
this manner, is shown in Figure 12. Two types of 
wire-resistance strain gauges have been used in this 
project. The first was the Baldwin-Southwark Type 
C-l, SR-4 gauge, in which the active wire is mounted 
on a thin paper backing that can be cemented to the 



Figure 12. Record of strain gauge mounted on the gun 
barrel. 


d Furnished by A. C. Keller of the Bell Telephone Lab¬ 
oratories. 


CONFIDENTIAL 























INSTRUMENTATION 


surface of the gun. For the purpose of measuring the 
expansion of the bore, these gauges were mounted so 
that the long dimension lay in a plane perpendicular 
to the axis of the bore. 

Gauges of the other type were made up directly on 
the gun, of 1-mil enameled Advance wire, one or two 
turns of which were wound around the gun and ce¬ 
mented to hold them in place. The principal advan¬ 
tage of this second type of gauge might be that its po¬ 
sition can be determined more precisely because it oc¬ 
cupies only 0.001 in. of the length of the gun, as com¬ 
pared with 0.16 in. for the C-l gauge. This was not a 
practical advantage, however, because the effects of 
the gun jacket on the strain records make it unneces¬ 
sary to know the linear position of a gauge more 
accurately than to 0.3 in. 

These strain gauges were used also for measure¬ 
ments of band pressure during rounds 61 through 77. 
Tangential and axial strains were measured at 90 in. 
and at 121.2 in. from the muzzle. Measurements of 
strain were made on the 37-mm gun. 

An important limitation of the strain-gauge meth¬ 
od, as it concerns the 3-in. gun, is the fact that im¬ 
portant parts of the gun tube are covered by the gun 
slide and the locking ring. The records obtained show 
that if it were possible to locate the strain gauges 6 in. 
apart all along the barrel, the signals from all of them 
could be recorded on one film. The connection of 
several strain gauges to a single oscillograph requires 
special design of the electric circuits. 39 

4,4,11 Solenoids for Projectile Velocity 
on Range 

The method used for determining the velocity of 
the magnetized projectile after ejection is a modifica¬ 
tion of the camera-chronograph method described in 
Army Ordnance Proof Manual. 2 ™ The changes consist 
in making the record with a cathode-ray oscillograph 
instead of a galvanometer, and in passing the tran¬ 
sient currents from the solenoids through suitable 
mixing circuits so that other phenomena can be re¬ 
corded on the same oscillograph trace. 

Each solenoid, of nominally 30 in. diameter, con¬ 
tained 7 layers, of about 35 turns per layer, of No. 26 
AWG cotton-enamel magnet wire, impregnated with 
insulating varnish and baked. Each solenoid was 
rigidly supported by a wooden frame mounted on one 
of the concrete piers of the range. In the earlier stages 
of the investigation four solenoids were used, their 
distances from the muzzle being 100, 200, 350, and 


91 

450 ft.; later, two more were added, one 50 ft, the 
other 500 ft from the muzzle. 

In a few of the later rounds two solenoids, 10 in. 
in diameter were used. With these solenoids the same 
deflection of the trace was obtained with the impor¬ 
tant advantage of considerably fewer turns and hence 
much less inductance. 

In Figure 5, the second curve from the bottom of 
(A) shows the form of the recorded signal produced by 
the passage of the magnetized projectile through a 
solenoid. 

4 412 Gauges for Powder Pressure 

Resistance gauges were usually used to measure 
powder pressure. In one form of these gauges the 
longitudinal compression of a piston by the powder 
pressure changes the resistance of a winding of fine 
wire cemented to the piston. This winding is prefer¬ 
ably applied so that the reflexed wires are parallel to 
the axis of the piston, as shown in Figure 13. In this 
type of gauge the effect of the pressure is to decrease 
the resistance of the winding, the relation between 
pressure and decrease of resistance being closely linear. 


Figure 13. Resistance gauge for powder pressure meas¬ 
urements. 




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92 


EXPERIMENTAL BALLISTIC FIRINGS 


Gauges with the winding applied as a helix around 
the piston are easier to construct but so much less 
sensitive that a preamplifier was required between 
the gauge and the oscillograph. In this type of gauge 
the resistance of the winding increases linearly with 
the applied pressure. 

In a third type of gauge, used for lower pressures 
(that near the muzzle and in the recoil cylinders), the 
pressure to be measured is applied to the inside of a 
thin-walled steel tube having an external helical wind¬ 
ing. These gauges were calibrated by applying a 
known load to the piston and measuring the change 
of resistance of the winding. 

In the early stages of the investigation some use 
was made of piezoelectric pressure gauges also. After 
round 19 it was found expedient to use the resistance 
gauge exclusively because of its greater ruggedness 
and convenience in calibration. The records obtained 
from the resistance-type gauges and the piezoelectric- 
type gauge may be compared by referring to Figure 
47 of NDRC Report No. A-229. 39 


4,413 Radiation Pyrometer for the Measure¬ 
ment of Powder-Gas Temperature 

The method of measuring the temperature of the 
powder gas is given in Section 2.5.5. Both the two- 
color pyrometers and single-color pyrometers de¬ 
scribed in Section 2.5.4 were used. The single-color 
pyrometers required less recording apparatus than 
the two-color. At least two oscillographs and their re¬ 
corders were required for each two-color pyrometer, 
whereas a single oscillograph and recorder sufficed for 
the single-color pyrometer. In any given round, the 
number of locations at which temperatures were re¬ 
corded was usually limited by the number of oscillo¬ 
graphs available. 

4414 Microwave Interferometer 

The motion of the projectile in the bore has been 
studied by means of a microwave interferometer, 108 
with which the distance from the projectile to the 



NOTE 

F?2 indicates the receiver r 2 

TURNED THROUGH 90° TO SHOW 
THE DIMENSION OF THE BOX 
PERPENDICULAR TO THE PAPER 



Figure 14. The radiator (Ri) and receiver (R 2 ) of the microwave interferometer, with their antennas (Ai and A 2 ), attached 
to the muzzle of a gun. 


CONFIDENTIAL 
























































































INSTRUMENTATION 


93 



Figure 15. Oscillograph trace of microwave interferometer. 


muzzle is measured continuously by reflected electro¬ 
magnetic waves several centimeters in wavelength. A 
beam of radiation is directed across the muzzle from 
a simple radiator, Ri, to an identical receiver, R 2 , 
shown in Figure 14. Ai and A 2 are antennas. Part of 
the beam enters the gun and travels down to the 
projectile; here it is reflected and returns to the muzzle, 
where it is picked up by the receiver. If at this point 
the reflected wave is in phase with the direct wave, 
the received signal is a maximum; if it is opposite in 
phase the signal is a minimum. Thus while the pro¬ 
jectile is moving along the bore the signal passes 
through a series of maxima and minima as the path of 
the reflected beam is shortened. The signal from the 
receiver is demodulated by a crystal detector and 
applied to an oscillograph. Figure 15 shows a sample 
of the resulting record. 

The calibration of the microwave interferometer 
was done by placing the projectile manually at defi¬ 
nite positions in the gun and reading the deflection of 
the galvanometer in the receiving circuit. This was 
done in steps of 0.01 or 0.02 ft from an arbitrary refer¬ 
ence point. The abscissa at each maximum and mini¬ 
mum of the oscillograph trace obtained in a firing 
gave the time of arrival of the projectile at a position 
which is given by the abscissa at the corresponding 
maximum or minimum of the appropriate calibration 
curve. Additional points were also obtained at inter¬ 
mediate points, especially near the start of motion. 
From the positions of the maxima and minima about 
70 values of position and of time were obtained. 

It is essential, both in calibrating and in firing, that 
the same sending wavelength and the same type of 
projectile be used. The shape and structure of the 
projectile nose modify the waveform of the detector 
current, the waveform being particularly influenced 
near the beginning of travel and as the projectile 
approaches the muzzle. 

The use of the microwave interferometer requires 
that the generator wavelength be within the limits 
2.62r and 3.42r, where r is the radius of the bore of 
the gun: 46 hence the wavelength used was nearly 10.5 
cm for the 3-in. gun and 5.0 cm for the 37-mm gun. 
Outside of these limits, more than one mode of vibra¬ 
tion appear, and the pattern is no longer simple. 


The interferometer method assumes that the posi¬ 
tions of maxima and minima are the same when the 
gun is fired as they were during the previous calibra¬ 
tion. However, when the gun is fired, the gas that is 
being pushed out of the bore by the projectile is high¬ 
ly compressed and may include some ionized powder 
gas that has leaked past the projectile. Consequently, 
the dielectric constant may differ significantly from 
that of empty space, and the positions of the maxima 
and minima may be shifted somewhat. The microwave- 
interferometer values were checked against barrel- 
contact values and the agreement was always within 
0.02 ft. 

4 415 Optical Ejection Indicator 

An optical system was arranged to send a beam of 
light perpendicular to the axis of the gun and to focus 
it on the axis a few millimeters in front of the muzzle. 
This beam was collected by a lens on the opposite 
side of the muzzle and illuminated the sensitive sur¬ 
face of a photoelectric cell. It was expected that, when 
the nose of the projectile cut the beam at the focus, 
the light would be cut off from the photoelectric cell, 
causing a deflection on the recording trace and indicat¬ 
ing the instant of ejection. However, it was found 
that smoke or other obscuring gases preceding the 
projectile gave a false indication of ejection. It was 
discovered in the records on the 37-mm gun that at 
the time the base of the projectile was ejected, the 
light from the hot gases produced a very sharp de¬ 
flection of the record from the photocell. 

4,416 High-Speed Cameras 

The smoke and flash were photographed by two 
types of cameras. One was a high-speed movie camera 
run at a speed of 3,000 pictures per second. The other 
camera was one in which the film was mounted on a 
rotating drum and which had a set of shutters of the 
focal-plane type under a row of lenses. A second shut¬ 
ter opened in synchronism with the firing of the gun. 
Successive pictures of a point were about 3.5 in. 
apart on the film and the time between pictures was 
approximately 2.5 msec. The movie camera made 


CONFIDENTIAL 





94 


EXPERIMENTAL BALLISTIC FIRINGS 


more pictures per second than the drum camera but 
had to be brought from rest to the correct speed at 
the correct time to record the desired event, which 
made synchronizing difficult. It also required intense 
illumination. The drum camera was more easily 
handled and did not require as intense illumination 
as the movie camera. 

4417 Flashmeter 

The flashmeter used to measure the intensity of the 
muzzle flash consisted of a radiation pyrometer mount¬ 
ed on the end of a 2-in. tube. The pyrometer is the 
same as used for the measurement of gaseous radia¬ 
tion from the powder chamber. The tube was pointed 
at the point where it was desired to measure the flash 
intensity. 

4,4 ’i8 "Microflash” Equipment 

The “Microflash” apparatus 6 and cameras were 
placed inside a small hut which had pipes about 14 in. 
in diameter extending 18 in. outside in the direction 
of fire. The projectile passed through the pipes and a 
microphone actuated by the bow wave of the projec¬ 
tile triggered the micro flash when the projectile was 
in the center of the hut. The shutters of the cameras 
were opened when the gun fired and remained open 
until the projectile had passed. 

4,419 Comparator for Distance 

Measurements on Films 

The data obtained in a firing are recorded in the 
form of a set of oscillograms from which the times of 
all events and the magnitudes of certain events must 
be determined. The time coordinate is determined by 
making distance measurements along the length of 
the film and interpolating between the timing pulses. 
Distance measurements made at right angles to the 
line of motion of the film give values of those quanti¬ 
ties for which there is a continuous record. Both of 
these kinds of measurement are made with a com¬ 
parator. 

The comparator consists of a heavy metal frame 
about 120 cm long, 30 cm wide, and 30 cm high, on 
which a traveling carriage rests at an angle of 45° 
from the horizontal. The carriage, driven by a screw, 


e Made by the General Radio Company, Cambridge, Massa¬ 
chusetts. 


has a range of about 90 cm longitudinal motion. Two 
long glass plates for holding the film, illuminated from 
below by a green fluorescent light, are set about 6 cm 
below the carriage and parallel to it. On the carriage 
itself are the observing microscope, supported on a 
Gaertner micrometer slide, and an auxiliary micro¬ 
scope moved by a micrometer screw. The slide has a 
range of 10 cm and a precision of 0.001 cm. All cross¬ 
wise distance measurements were made with it. 

Running along the top of the comparator frame is 
a scale, 100 cm in length, divided in millimeters. It is 
not a precision scale, but is more than accurate enough 
to give the position of the carriage to the nearest 
millimeter. Below the carriage, in a position such that 
the auxiliary microscope on the carriage can be focused 
on it, is an Invar meter bar made by the Geneva 
Society, divided in millimeters and accurate to within 
0.0003 mm at 20 C. The two scales are rectified by 
setting the carriage approximately on a millimeter 
division mark on the top scale, setting the cross hair 
of the auxiliary microscope on a millimeter division of 
the meter bar and then, while holding the micrometer 
screw fixed, turning the micrometer scale (which is 
held by a friction clamp) until the zero mark of this 
scale coincides with its fiducial mark. The scales re¬ 
main sufficiently well in alignment over the entire 
length of the film so that there is no uncertainty. 

The micrometer screw attached to the auxiliary 
microscope has a pitch of 1 mm, and its head is divid¬ 
ed into 100 divisions. Thus the longitudinal position 
of the observing microscope is given to 1 mm by the 
top scale and to 0.01 mm by the reading of the microm¬ 
eter screw when the cross hair of the auxiliary micro¬ 
scope is in coincidence with a millimeter division of 
the meter bar. Although fractions of a hundredth of a 
millimeter could be estimated on the micrometer 
screw, the error of setting the cross hair on the divi¬ 
sions of the scale may be as much as 0.01 mm. 

4 - 4 - 20 Protractor for Differentiating Curves 

The instrument shown in Figure 16 furnishes a rap¬ 
id and simple method for graphical differentiation. 
A steel protractor, divided in degrees with a vernier 
reading to 5 min, is equipped with a gear for slow 
motion. It is attached to the straightedge of a draft¬ 
ing machine. The sheet carrying the curve to be dif¬ 
ferentiated is placed on the drawing board so that the 
axis of abscissas is parallel to the straightedge. The cen¬ 
ter of the protractor circle is placed at the point where 
the slope is required. The protractor is then turned 


CONFIDENTIAL 





PROJECTILE GAUGES 


95 


until its edge is tangent to the curve at that point, 
and the angle is read. The derivative at that point 
is directly proportional to the tangent of the angle. 

45 PROJECTILE GAUGES FOR INTERIOR 
BALLISTIC MEASUREMENTS 

4,5,1 Introduction 

In addition to the comprehensive ballistic investi¬ 
gations at Carderock, described in the earlier sections 
of this chapter, Division 1 sponsored the develop¬ 
ment of techniques and instruments for making 
ballistic measurements directly on a projectile mov¬ 
ing in the bore of a gun. The quantities to be measur¬ 
ed were displacement, velocity, and acceleration of 


the projectile, pressure on the base of the projectile, 
and friction between the projectile and the barrel. 

The basic principles 5 involved had been developed 
by Section T of Division A, NDRC. Then the Geo¬ 
physical Laboratory of the Carnegie Institution of 
Washington undertook to perfect for Division 1 an 
experimental routine for use with different sizes of 
projectiles. 46 

4 5 2 Wire-in-Bore Technique f 

The basic technique for these measurements was 
that of making contact with the moving projectile by 


f A similar technique had been developed by the British at 
about the same time. 337 



Figure 16. Protractor modified for differentiation of a curve. 


CONFIDENTIAL 





































































96 


EXPERIMENTAL BALLISTIC FIRINGS 


means of a wire stretched coaxially within the bore 
of the gun from the nose of the projectile to the muz¬ 
zle and picked up by a “collector cup” on the nose of 
the projectile. By making electrical contact with the 
projectile in this way, many types of measurement 
are possible, for any phenomenon associated with a 
projectile that can be converted into an electrical sig¬ 
nal may be measured. 

Projectile Displacement 

The simplest application of this technique was to 
measure the displacement of a moving projectile by 
measuring the change of resistance of the wire as its 
length became less. This was done by recording with 
an oscillograph the variation of the voltage across the 
ends of the wire while a small constant current was 
passed through it. A lead lining in the collector on the 
nose of the projectile was found to be necessary for 
maintaining good electrical contact. The oscillo¬ 
graphic trace showed fluctuations in the resistance of 
the contact between the projectile and the bore sur¬ 
face before the rotating band was engraved. 

Improved Obturation 

An apparent decrease in the resistance of the wire 
at the beginning of motion of the projectile was 
attributed to heating of the wire by powder gas that 
had leaked past the projectile. Several methods were 
undertaken to improve obturation. One method pro¬ 
vided, to a degree, an initial seal which was effective 
only until the band was engraved. It consisted in 
spinning up one or more skirts of metal from the band 
with a cutting tool. Thus modified, the projectile had 
to be forced home in loading. The spun skirts gave no 
added obturation after the band passed through the 
forcing cone. The other method for improving obtura¬ 
tion was the use of the Bridgman “unsupported-area 
pressure seal,” referred to as a “hydraulic obturator.” 
Various combinations of the two methods of obtura¬ 
tion were tried; the most effective was found to be 
the hydraulic method in combination with two skirts 
spun up on the band. 

With this best method of obturation, various kinds 
and sizes of wires were tried, in particular those kinds 
having a low temperature coefficient of resistance. 
It was found that No. 34 manganin wire gave the 
best-appearing traces. The magnitude of the overall 
error in measuring the displacement in the 20-mm 
gun was estimated as + 3.2 cm. For comparison, 


measurements were made using the microwave tech¬ 
nique, 8 for which the error was found to be less than 
one-third as large. 

Piezoelectric Crystal Gauges 

The wire-in-bore technique described in the pre¬ 
vious section made it possible to obtain a signal 
from a piezoelectric crystal gauge located in a pro¬ 
jectile and subjected to various forces associated with 
the motion of the projectile in the bore of a gun. Such 
gauges were developed for the measurement of accel¬ 
eration, base pressure, and bore friction. 5 

The chief difficulty in making these gauges was in 
designing the component parts small enough to fit 
into 20-mm projectiles and yet strong enough to with¬ 
stand the high acceleration (100,000 g) to which they 
were to be subjected during firing. Thus in order to 
reduce crystal breakage it was possible to use only one 
crystal, although a stack of three would have been 
preferable as far as the strength of the signal was con¬ 
cerned. Mechanical refinements had to be introduced, 
such as ball-and-socket joints to equalize the pressure 
over the surface of the crystal. 

The general qualitative results obtained by the use 
of the crystal gauges may be briefly stated as follows. 
The signal traces obtained from the three types of 
gauges were consistent in shape during the first one- 
third of the travel time. The last part of the trace was 
anomalous in all cases; furthermore, the anomaly was 
not consistent in size or in shape. Special experiments 
showed that this part of the signal was associated 
with the leakage of ionized propellant gas past the 
projectile as it traveled through the bore. No effective 
way was found to reduce to an insignificant amount 
the spurious signals from this cause. For the first 
fourth of the travel of a 20-mm projectile the maxi¬ 
mum error of the acceleration gauge amounted to 
± 10 per cent, that of the base-pressure gauge ± 5 to 10 
per cent, and that of the bore-friction gauge ± 10 to 
20 per cent for starting friction and + 40 to 60 per 
cent for running friction. 

Preliminary experiments 46 with some 3-in. projec¬ 
tiles fitted with crystal gauges and fired from the 

g The microwave interferometer described in Section 4.4.14 
was an outgrowth of an apparatus developed as part of this 
project for measurement of the displacement of a projectile in 
the bore of a 20-mm gun. 46 After the possibility of such a 
measurement had been suggested by Dr. T. H. Johnson of the 
Ballistic Research Laboratory, Aberdeen Proving Ground, 
parallel development of apparatus for the purpose was carried 
out there and by Division 1 at the Geophysical Laboratory. 


CONFIDENTIAL 





PROJECTILE GAUGES 


97 


3-in. gun at Carderock showed that the leakage of 
powder gas past the projectile was equally as serious 
as in the 20-mm gun. The record was unreliable dur¬ 
ing the last 4 msec of travel time, that is, after the 
projectile had traveled 50 cm, which is only about 
one-seventh the length of travel. During the first part 


of the travel the maximum uncertainty was estimated 
to be of the same order of magnitude as for the 20-mm 
gun. Most of this uncertainty was caused by gas 
leakage. Thus these gauges are not satisfactory un¬ 
less some more effective means of obturation can be 
found. 


CONFIDENTIAL 



Chapter 5 

HEATING OF GUNS DURING FIRING 

By H. L. Black a and G. Comenetz h 


51 INTRODUCTION 

I n attacking the problem of hypervelocity, the 
direct approach would seem to be that of increasing 
the energy of the powder charge. This could be done 
quantitatively by increasing the amount of powder or 
qualitatively by using a powder with higherpotential. 
The limitations of the former method are well known. 
When the latter method is used we encounter higher 
flame temperatures, and the resulting temperature of 
the gun bore surface becomes seriously high. If we 
combine with this increased flame temperature the 
high rates of fire maintained in automatic weapons, 
the question of heating of guns during firing is critical. 
The thermal effects are an important factor in erosion 
and in addition lower the tensile strength of the gun 
metal. Division 1 was therefore interested in the 
study of the heat input to the gun, the distribution of 
heat in the barrel, the temperature of the bore surface, 
and in methods of gun cooling. 

As a result of these studies, our knowledge concern¬ 
ing heat input from a single round, input from a burst 
or a series of bursts, temperature attained by the bore 
surface, critical temperatures in firing, and methods 
of cooling has been greatly extended. The application 
of this knowledge in the development of improved 
machine gun barrels by lining or plating them has 
been of definite value and will aid in the design of 
future hypervelocity guns. In particular, these tem¬ 
perature investigations show that to combat the 
effects of heat input we must look to the develop¬ 
ment of alloys with physical qualities which will make 
them satisfactory as liners or coatings for the bore 
surface at temperatures above the melting point of 
steel. The search for such alloys is reviewed in Chap¬ 
ter 16. 

A better understanding of the heat input from 
band friction and its combination with that from 
powder gases is to be desired. Similarly, the effect of 


“Technical Aide, Division 1, NDRC. (Present address: De¬ 
partment of Mathematics, Michigan State College, East Lan¬ 
sing, Michigan.) 

b Mathematical Physicist, Geophysical Laboratory, Carne¬ 
gie Institution of Washington. (Present address: Westinghouse 
Research Laboratories, East Pittsburgh, Pennsylvania.) 


thermal stresses taken in combination with band and 
powder pressures should be more completely analyzed 
as an aid to the designer. We may expect improved 
experimental methods of measurement of the bore 
surface temperature. Finally, methods of cooling 
un der service conditions certai nly n eed to be improved. 

52 HEAT INPUT TO THE BORE SURFACE 
Magnitude of the Energy Loss 

A complete analysis of the distribution of the 
energy of the charge is given in Section 6.5. Here we 
are particularly concerned with the percentage of the 
energy expended in heating the gun; for purposes of 
comparison, however, it is convenient to compare this 
with the kinetic energy of the projectile at the muz¬ 
zle. It is customary to consider heat transferred to the 
cartridge case as going to the barrel. 

Experiments 516 as early as 1870, in which rifles us¬ 
ing black powder were tested, indicated that from 
31.7 to 39.3 per cent of the total energy of the powder 
was absorbed by the barrel. A later investigation 442 
with guns gave a heat input ranging from 4.2 to 17 per 
cent of the powder potential. The energy balance for 
one shot fired in a rifle has been determined in a series 
of carefully controlled experiments. 450 

Since 1936 a number of measurements and calcula¬ 
tions of energy distribution have been made. Table 1 
presents the results obtained by a number of experi¬ 
menters for various guns. In Tables 2 and 3 are given 
calculated values for typical guns in a similar range 
of calibers. 

These tabular percentages do not, of course, repre¬ 
sent results which are completely comparable. The 
powders used vary; however, for the larger guns, all 
powders have flame temperatures close to 2500 K. In 
Section 15.5.3 it is shown that heat input is a linear 
function of the flame temperature of the powder in a 
caliber .50 gun. 

The data available are too limited to draw a valid 
conclusion as to the functional relationship between 
caliber and heat input; it is, however, apparent that 
as caliber increases, the percentage of the powder po¬ 
tential transformed into heat decreases. This is in line 


98 


CONFIDENTIAL 



HEAT INPUT TO THE BORE SURFACE ^9 


Table 1 . Percentage of energy of powder lost by heating of gun compared to that acquired by the projectile measured 

for various guns. 






Cal .30 

Cal .50 MG 182183 

3-in. A A, M3 179 

3-in. 


rifle 450 

(1) (2) 

(1) (2) 

Navy 65 gun 

K. E. of projectile 

33 

29 28 

29 32 

27.7 

Heating of gun 

22 

19 16 

11 9 

6.3 


Note. Caliber .50 MG: (1) water-cooled barrel fired without water-cooling jacket; (2) air-cooled heavy barrel. 
Three-inch AA, M3: (1) with Pyro powder; (2) with NH powder. 


Table 2.* Percentage of energy powder lost by heating 
of gun compared to that acquired by the projectile— 
calculated for U. S. guns. 48 


Gun < 

Cal .50 

37-mm 

3-in. 

4.7-in. 

8-in. 

16-in. 

K.E. of projectile 

29 

32 

23 

27 

25 

26 

Heating of gun 

18 

14 

8 

8 

7 

5 


* Compare Table 2 of Chapter 3 which gives for these same guns corre¬ 
sponding values for the heat loss just to the time of ejection. 


Table 3. * Percentage of energy of powder lost by heat¬ 
ing of gun compared to that acquired by the projectile— 
calculated for British guns. 348 



2-pr. 

25-pr. 

8-in. 


(1.57 in.) 

(3.65 in.) 

Mk. VIII 

Gun 

S.C. Cordite, 

# Cordite 

Cordite 

Propellant 

slotted tube 

W .057 

S.C. 205 

Travel (calibers) 

36.67 55.01 73.32 

22.06 

43.32 

K.E. of projectile 

20.7 26.9 31.1 

41.8 

40.3 

Heating of gun 

4.7 6.5 8.1 

4.6 

4.7 


* The method of calculation is that described in Section 5.4.1 under 
“Method of Hicks, Thornhill, and Ware.” 


with the theory that most of the heat is transferred 
from a laminar layer of gases relatively close to the 
bore surface, rather than from the main gas stream, 
which occupies a larger percentage of the total volume 
in larger guns. 

5,2 2 Sources of Heat Input 

Studies of the causes of heat input to the bore sur¬ 
face are numerous in ballistic literature. 144 348 ’ 505 The 
sources may conveniently be classified as follows: 
(1) radiation from the hot powder gases; (2) input 
from gas leakage past the rotating band; (3) heat of 
engraving and friction; (4) transfer from the gases in 
a laminar layer near the bore surface. 

Radiation 

Radiation is an unimportant source of heat input. 
In the heat transfer equation, 

Q = (r e l€2 (7V - 7V). (1) 


it has been shown 12 that for a gas temperature of 
2500 K above that of the surface, the maximum rate 
of transfer is 53 cal per sq cm per sec. T 0 is the gas 
temperature, T s the temperature of the bore surface, 
6i and e 2 are the emissivities, and a has the value 
1.36 X10 -12 cal per sq cm per sec per fourth power of 
the degree of temperature. At this rate the bore sur¬ 
face temperature would be raised only about 36 
centigrade degrees during the firing of the very largest 
guns. 

Heat of Engraving and Friction 

There has been a great deal of argument as to the 
importance of mechanical effects as sources of heat 
input. 0 Thus as recently as 1925 it was suggested on 
the basis of extensive experimentation “that the 
main source of heating of the bore does not lie in the 
transfer of heat from the powder gas to the wall of the 
barrel but in the mechanical work (friction, engrav¬ 
ing by the rifling, and so on) which is converted into 
heat.” This, however, was a minority opinion, and the 
review continued: “It is to be hoped that still further 
investigations may be carried through in order to 
solve the interesting question as to the causes of the 
heating of the bore.” 

With this in mind, it was decided by Division 1 to 
make experimental determinations of the heat input 
from engraving and friction using new methods of 
measurement. 72 Because of the interest in pre¬ 
engraved projectiles (Section 27.3), a logical first 
approach seemed to be by a comparison of the 
input with ordinary and with pre-engraved pro¬ 
jectiles under otherwise identical conditions. 

First measurements were made by the calorimetric 
ring method described in Section 5.4.3, and an early 
report stated, “Preliminary results show only little 
variation in the difference of heat input values for 
standard and pre-engraved M2 bullets measured at 


c Opinions of some of the writers were summarized by 
Cranz.’ 144 * 605 


CONFIDENTIAL 

















100 


HEATING OF GUNS DURING FIRING 


the forcing cone against values of that difference 
determined in back of the origin of rifling, thus indi¬ 
cating that the heat of engraving is rather small com¬ 
pared with the total heat input to the bore.” Later 
measurements during firings supported this position, 
but it was considered advisable to attempt a direct 
determination of the heat of engraving and friction. 

For this purpose a falling-weight apparatus was 
devised. A heavy carriage dropped from a maximum 
height of 64 ft struck a piston which in turn forced a 
bullet through an “engraving ring.” This was a short 
section of a caliber .50 barrel placed vertically and 
supported by a hollow metal cylinder. The ring acted 
as a calorimeter; from its rise in temperature the heat 
input from friction and engraving was calculated. 
Formulas due to Jaeger were used to extrapolate to 
velocities of the order of those attained in actual 
firings. The results indicated an input for ball bullets, 
M2 of approximately 1.3 cal/cm 2 near the beginning 
of travel, decreasing to 0.6 cal/cm 2 toward the muzzle. 
This value compares with a total heat input during 
firing of the order of 10 cal/cm 2 . 

For reasons which; will be discussed in Section 5.4.4, 
this input does not, however, cause a very large varia¬ 
tion in the maximum bore surface temperature. This 
seeming contradiction explains the varying estimates 
of the contributions of engraving and friction to heat 
input. One such estimate 397 of their importance has 
been given as follows: “During the engraving of the 
driving band, a short section of the barrel surface 
near the commencement of rifling may, at worst, be 
raised to the melting temperature of the band for a 
small time interval. Temperature calculations are 
made, including a liberal estimate of this heat trans¬ 
fer, but it is not found to produce sudden large in¬ 
creases to temperatures above the melting point of 
the barrel, and in the 8-in. gun gives an increase of 
only 200 centigrade degrees in the maximum surface 
temperature at the commencement of rifling. This 
estimate is probably several times too great.” This 
question of the effect of input from the rotating band, 
in combination with input from the gases, is discussed 
in Section 5.4.4. 

Frictional heat input will continue even after en¬ 
graving is considered complete. It is generally agreed, 
however, that this is relatively unimportant and that 
it causes a negligible increase in maximum bore sur¬ 
face temperatures. Friction does, however, add ap¬ 
preciably to the strains upon the gun. 39 - 65 ’ 113 It is also 
important to consider the melting of the metal of the 
rotating band, 556 which causes coppering of the bore. 


Even here the effect is not so great as might be ex¬ 
pected because the higher diffusivity of the common 
band materials tends to more than counterbalance 
the lower melting point. This is brought out in the 
theoretical discussion in Section 6.2.2. 

By assuming that the rotating band in a 37-mm 
gun, M3 rises to a maximum temperature 1000 centi¬ 
grade degrees above the temperature of engraving, a 
coefficient of friction of 0.03 was calculated. 45 Based 
on this value, curves for rates of heat generation and 
conduction were plotted against time, and for heat 
transferred to the bore surface as a function of posi¬ 
tion in the bore. It was concluded that conduction 
into the bore reached a maximum slightly above 2 
cal/cm 2 . 

Transfer from the Gases in a Laminar Layer 
Near the Surface 11 

This brings us to the consideration of the source of 
input now generally considered the most important, 
transfer from the hot powder gases. The powder gases 
rush past the bore surface at a velocity of the order of 
1,000 fps, and heat is transferred by forced convection. 
For the gas flow we may use the theory for an incom¬ 
pressible fluid; this is justified on the basis of experi¬ 
ments 496 that showed that the shearing stress on the 
wall of a pipe (or gun) is the same for compressible as 
for incompressible fluids at the same Reynolds num¬ 
bers. 

If now we consider the motion of the gases as turbu¬ 
lent, it nevertheless is true that the velocity near the 
walls remains laminar, although with a large gradient. 
At some distance 8 from the walls, the velocity ap¬ 
proaches its midstream value; and although the value 
of 8 is not sharply defined, this laminar layer is re¬ 
ferred to as a film. The film thickness and, hence, the 
rate of heat transfer depend on the condition of the 
bore surface. 

The heat transfer coefficient h may be defined by, 

Q = h (T g — T s ) (2) 

in which Q is the actual rate of heat transfer, while T a 
and T s are, respectively, the temperature of the pow¬ 
der gas (beyond the laminar layer) and of the bore 
surface. The aim of the hydrodynamic analysis is to 


d A standard work on heat transfer is that of McAdams; 613 
for related hydrodynamic problems, Goldstein 607 may be con¬ 
sulted. Fundamental theory discussed in these works is referred 
to in NDRC Reports A-87, A-201, and A-262. The following 
discussion is based largely on these three reports. 


CONFIDENTIAL 




HEAT INPUT TO THE BORE SURFACE 


101 


evaluate h in terms of the state of the gas and the con¬ 
dition of the bore surface. Because it is difficult to 
specify the latter precisely, that is, to define the de¬ 
gree of roughness of the particular wall, no purely 
theoretical formula can be given with perfect confi¬ 
dence. A productive procedure, 48 however, has been 
to obtain an expression for h with one adjustable con¬ 
stant that is related to the roughness of the wall; and 
to calculate from the observed heat transfer at var¬ 
ious positions in various guns an empirical value of 
this constant that might be applicable to the bore 
surface of all guns. This procedure has proved fairly 
successful, because the heat transfer coefficient is not 
very sensitive to minor variations in roughness. Meas¬ 
urements of the friction when water flowed through 
rifled and unrifled caliber .50 barrels showed that the 
presence of the rifling made only a negligible differ¬ 
ence. 165 

The fundamental hydrodynamic expression for h 
(expressed in calories per square centimeter per sec¬ 
ond) is given by 

h = \\c p pU (3) 

where c p is the specific heat, p the density, U the ve¬ 
locity of the powder gas, and X is an empirical con¬ 
stant, the friction factor or Fanning coefficient. X is 
related to the bore diameter and to a roughness para¬ 
meter r, which measures the average size of the irreg¬ 
ularities on the bore surface. Different expressions of 
this relationship are required if the surface is hydro- 
dynamically “smooth” (r < 10~ 5 cm), or “rough” 
(r > 10 -5 cm).'The latter condition almost certainly 
applies in guns. 

The application of equation (3) to the calculation 
of bore surface temperatures and heat input is dis¬ 
cussed in some detail in Section 5.4.1. The preferred 
analysis 48 of the most careful measurements 71 of heat 
transfer near the origin of bore of a caliber .50 barrel 
leads to the value of X given by 

X = (14.2 + 4 log 10 D)~ 2 (4) 

where D is the caliber of the gun in centimeters. 

In some computations, described in Section 5.4.1, 
equation (3) has been simplified. Using an average 
ballistic value for c p , and expressing p in the con¬ 
venient units of grams per cubic centimeter, and U' 
in 1,000 fps, we have for rough pipes, 

h = 10,400 \ P U'. (5) 

For smooth pipes, h may be evaluated in terms of 
pU'. Tables of the dependence have been given, 12 and 


the results may be approximated to yield 

h = 8.9 (pU) 0 - 8 . (6) 

These last two equations are probably not so ade¬ 
quate as the empirical expression (4), although the 
results to which they lead, as described in Section 
5.4.1, do not differ greatly. 

Most of the calculations of bore surface temperature 
and of heat input, including the one from which 
equation (4) was derived, have been based on the as¬ 
sumption that input from sources other than transfer 
from the gases was relatively unimportant. Actual 
input as determined by measurements by thermo¬ 
couples (Section 5.3) give approximate figures for the 
transfer at various positions along the barrel and 
therefore may be used to check the assumptions. It is 
perhaps significant that in the cases of the 37-mm 
gun, T47 and the 3-in. gun at Carderock, described 
in Chapter 4, the experimental results show a slightly 
greater decrease toward the muzzle than was calcu¬ 
lated by the methods of NDRC report A-262. 48 

For example, in the 3-in. gun, near the origin of 
rifling, the experimental results 106 - 113 indicate an in¬ 
put, using two different methods, of 25.8 and 26.7 
cal/cm 2 , whereas the calculated input was 24.6 cal/ 
cm 2 . Nearer the muzzle, however, the calculated in¬ 
puts were from 20 to 30 per cent higher than those 
observed. In the case of the 37-mm gun, T47, using 
FNH-M5 powder and firing 1.62-lb projectiles with a 
muzzle velocity of 3,600 fps, the measured input near 
the origin averaged 23.5 cal/cm 2 as compared with a 
calculated value of 24.6; whereas with the cooler Ml 
powder the actual results were 16.4 cal/cm 2 as 
against a calculated value of 14.1. Again, nearer the 
muzzle the observed values fell below those calculated. 

These results constitute good confirmation of the 
general methods employed in calculating the heat in¬ 
put to the bore surface. The divergences between 
theory and experiment may be due to several causes. 
Neglect of the frictional contributions to the heating, 
both in the standard gun from which equation (4) 
was derived and in other guns, would lead to diver¬ 
gences of the order of 1 to 2 cal/cm 2 , if the frictional 
heat varies relative to the total heat; the ratio of 
frictional to total heat is certainly greater near the 
origin. The situation is further confused by the fact 
that the theoretical basis of Nordheim’s tables in¬ 
volves the total heat transfer f rom the gases, whereas 
the empirical calibration factor that is used to derive 
equation (4) involves the total transfer from the 
gases, friction and all other causes. 


CONFIDENTIAL 



102 


HEATING OF GUNS DURING FIRING 


Variations in surface roughness sufficient to cause 
variations in heat input may occur between guns, and 
even at different points along the bore of the same 
gun. Finally the actual ballistics may differ from 
those assumed in making the calculations; for ex¬ 
ample, the theory takes no account of a possible 
temperature gradient in the powder gas down the 
barrel, such as has been observed (cf. Section 6.5.1). 
It is clear that although the fundamental concept that 
heat transfer is principally due to forced convection 
is soundly based, a much greater body of experi¬ 
mental data is necessary before we can specify the 
reliability of the results to a degree which will enable 
us to separate the effects of the various factors. 

53 METHODS OF MEASUREMENT OF 
HEAT INPUT 

Overall Calorimetric Methods 

The early attempts to measure heat input were 
largely of a direct calorimetric type. One investi¬ 
gator 516 measured the rise in temperature of mercury 
with which the bore of a rifle was filled; from this rise 
the heat input was calculated. In a later experiment 442 
the hot gun barrel was put in a water bath. 

This method was also employed in a more precise 
determination described 144 - 505 as follows: “The barrel 
of the gun . . . was surrounded by a thin-walled 
cylindrical sheet-metal jacket of about 4.5-cm dia¬ 
meter, extending from the muzzle almost to the 
breech . . . during the firing, gun and metal jacket 
were enclosed in a wooden box. The jacket was filled 
with 1.2 liters of water, so that the entire barrel was 
in a water bath. A stirring ring was moved up and 
down in the water. The water equivalent [thermal 
capacity] of the whole arrangement was obtained by 
passing steam through it.” 

It has been pointed out 144 - 505 that a rough determi¬ 
nation of input is possible with the aid of a machine 
gun having a cooling jacket; but the hypothetical ex¬ 
ample given was based on the assumption of constant 
input per round. This method was applied 475 to a 
Maxim caliber .30 q.f. rifle, with the following re¬ 
sults: “If this quick firing rifle be originally at about 
60°F it will boil water after firing about 600 rapid 
rounds in about min.; it will then continue to 
evaporate water at the rate of pints per 1000 
rapid rounds [about 2.2 minutes].” From these data 
a heat transmission of 133,363 Btu per square foot 
of outer surface per hour was computed and it was 


concluded finally that the heat absorbed was nearly 
twice the muzzle energy. The corresponding estimate 
of the maximum bore surface temperature, using 
Fourier analysis, was only 600 F. 

This method may be applied to obtain the approxi¬ 
mate input in a water-cooled 40-mm Navy gun on 
which temperature measurements were taken. Inlet 
and outlet water temperatures were read for bursts of 
50 and 100 rounds under almost identical conditions, 
and for a 150-round burst with slight variation. Con¬ 
sistent results were obtained in the three cases. Using 
the first as an illustration, the burst was fired in 
25 sec; a peak temperature 25 centigrade degrees 
above the original was attained almost immediately 
afterward. Equilibrium inlet and outlet temperatures 
were reached approximately 3J^ min later at a value 
14 centigrade degrees above the initial value. The net 
input per round after this period of cooling is 6.0 
cal/cm 2 of bore surface; if it is assumed that all the 
water reached maximum temperature, the average 
input per round is 10.8 cal/cm 2 . It should be re¬ 
marked that the reduction of input with successive 
rounds is not so great as in a gun without water¬ 
cooling. 

In studies 182 - 183 at Aberdeen Proving Ground of the 
heat transferred to the cartridge cases of caliber .50 
machine guns, both air-cocled and water-cooled, the 
temperature rise of a quantity of water into which the 
cases were dropped during the firing was measured by 
means of a mercury thermometer. 

5 3 2 Artificial Heating of the Barrel 

A natural approach to the study of heat input dur¬ 
ing firing is to reproduce by other methods the tem¬ 
perature distributions that have been observed dur¬ 
ing firings. Such an investigation 450 was carried out 
using a rifle barrel itself as a resistance thermometer. 
A shot was fired every 30 sec until a “quasi-stationary 
state” was reached, with an average barrel tempera¬ 
ture of 122 C. In the second part of the experiment, 
an insulated manganin wire was passed through the 
barrel, and then heated by current from a storage 
battery until an equivalent state was reached by the 
barrel. The input from the wire was calculated, and 
from this the heat introduced by one shot was com¬ 
puted to be 624 calories. This represented an average 
input to the hot barrel. 

A refinement of this method has been reported by 
British investigators. 398 Rounds are fired at a rate 
such that a steady external temperature distribution 


CONFIDENTIAL 




METHODS OF MEASUREMENT OF HEAT INPUT 


103 




o 

UJ 

u. 

o 

UJ 

CL 

to 


u 

52 


o 


<0 

co 

UJ 


</> 


X 

o 


to 

z 

o 

to 

z 

UJ 


Q 


-J 

-J 

< 



CONFIDENTIAL 


Figure 1. Calorimetric rings in gun. (This figure has appeared as Figure 1 in NDRC Report No. A-399.) 



















































































































104 


HEATING OF GUNS DURING FIRING 


along the barrel surface is established and maintained; 
this distribution is recorded by means of thermo¬ 
couples. An electrical heating coil constructed of a 
spiral of resistance tape on a refractory form is then 
inserted into the barrel. The spacing of the turns of 
the tape is adjusted until the same steady external 
temperatures are obtained. The electrical energy dis¬ 
sipated is measured. From these data, the heat trans¬ 
fer per unit area of the bore surface per round can be 
deduced. It should again be noted that this will give 
results for transfer to a “hot” barrel. The method is 
not satisfactory for points near the origin of rifling. 


The problem, for a perfectly insulated ring, has also 
been approached by calculating the amount of heat 
stored in the bakelite gaskets. 

Consistent results were obtained by the calori¬ 
metric ring method. It gives absolute heat input de¬ 
terminations near the origin of rifling, where some 
other methods fail. Objections to it as compared with 
methods developed later are the large amount of shop 
work involved, the lower degree of precision, and de¬ 
terioration of the insulation resulting from firing. 

5 3 4 Calorimetric Section Method 71 


Calorimetric Ring Method 71 


Several methods of heat input measurement have 
been developed for Division 1 by the Leeds and 
Northrup Company. In the first of these a thermally 
and electrically insulated ring, cut from a barrel of 
the same caliber as the gun being tested, is set flush 
with the inside of the bore, as shown in Figure 1. The 
ring must be thick enough to insure against mechan¬ 
ical deformation. A thermocouple is affixed to its outer 
wall, with wires leading to the outside of the barrel 
through thin bakelite tubes. These are sealed against 
gas leakage by neoprene washers compressed by 
metal plugs. Fluorocarbon is used as an obturator; 
insulating gaskets made of linen-base bakelite are 
subjected to a make-up force of 20 tons during in¬ 
stallation. 

The first rings were placed behind the origin of 
rifling; later, rifled rings were installed at locations 
farther forward. These were carefully aligned with 
the rifling of the bore and pinned to prevent rotation. 

The heat input per unit area of the bore is com¬ 
puted from the formula 


Emc 

tv dwP 


(7) 


in which E is the emf output of the thermocouple, P 
its thermoelectric power, and m, c, d, and w respec¬ 
tively the mass, specific heat, inner diameter, and 
width of the ring. 

Radial thermal equilibrium is reached throughout 
the ring in less than a second. Heat is lost relatively 
fast by leakage, particularly through the insulating 
gaskets. The emf-time curve of the thermocouple 
reaches a maximum and then falls steadily. To obtain 
the ring temperature for zero thermal loss, the expon¬ 
ential decay curve can conveniently be extrapolated 
back to time zero on semilogarithmic graph paper. 


The term ‘‘calorimetric section” as used here applies 
to a relatively long section of the barrel which has 
been reduced in diameter by removal of an annular 
ring of metal. The purpose of this machining is to de¬ 
lay and diminish the longitudinal flow of heat to the 
remaining “thick” sections. As long as the thickness 
of the barrel wall of the calorimetric section is small 
as compared with the distance from the measuring 
thermocouple to the nearest end of the section, the 
section acts as a calorimeter for a considerable length 
of time after the heat input through the bore has 
taken place, and the temperature of the outer surface 
remains at its maximum value for about a second. 
Since this temperature is determined only by the 
amount of heat transferred through the bore surface 
and the thermal capacity of the section, measure¬ 
ments by this method give absolute values of the heat 
input. 

Since the heat input through the bore is equal to 
the heat received by the barrel calorimeter, we have 

Hwd = ecp-j [(<Z + 26) 2 - d 2 ] (8) 

where Pis the heat input per unit area, 9 the tempera¬ 
ture rise of the section, d the average inside bore di¬ 
ameter, b the wall thickness of the section, c the 
specific heat of the gun metal, and p its density. Solv¬ 
ing this equation for H , we obtain 

H = f(bcp) (l + (9) 

in which the symbols E and P have the same mean¬ 
ing as in the preceding section. 


5,35 Outer Surface Measurements 

Thermocouples attached to the outer surfaces of 
guns have been used in many temperature measure- 


CONFIDENTIAL 





METHODS OF MEASUREMENT OF HEAT INPUT 


105 


EMBEDDED THERMOCOUPLE 
27.3 IN. FROM MUZZLE 


EMBEDDED THERMOCOUPLE 



Figure 2. Assembly for embedded thermocouple. (This figure has appeared as Figure 2 in NDRC, Report No. A-434.). 


ments. As examples, the British report using the 
method in temperature measurements with 20-mm, 
Mark V (Hispano) guns, 3.7-in. Mark VI AA guns, 
and 3-in. and 4.2-in. mortars. 

The method has also been extensively employed in 
studies with the caliber .50 machine gun, especially 
by North American Aviation, Inc., where a large 
number of heating and cooling curves have been ob¬ 
served. 526 Since this type of measurement can be 
made under conditions very similar to service ones, 
there has been great interest in the study of the cor¬ 
relation between outer surface temperature and mal¬ 
functioning in machine guns (Section 5.6.4). 

The unmodified barrel wall itself may be considered 
as a calorimetric section in the sense of Section 5.3.4. 
Although this method is a convenient one for absolute 
heat input measurements in large caliber guns, it has 
some disadvantages. 71 The accuracy of measurement 
is reduced because of the greatly diminished tempera¬ 
ture rise in a thick-wall calorimeter; the equilibrium 
conditions are attained slowly; and the. temperature 
indication of the thermocouple is averaged over a 
large bore area. This method was used 106 in conjunc¬ 


tion with the ballistic firings of the 3-in. and 37-mm 
guns at Carderock (Section 4.2). 

5 3 6 Embedded Thermocouple 

At a distance of more than a few hundredths of an 
inch from the bore surface, the transient rise of tem¬ 
perature as a function of time when a single round is 
fired depends on the shape and ’material of the barrel 
and the. total heat input front the round, but not 
sensibly on the rate of heat input as a function of 
time, because the time during which the heat enters 
is so short. If the distance from the bore’ is several 
times larger than the depth'of rifling and at least 
several times smaller than the distance back to the 
origin of bore or forward to the muzzle, the time-tem¬ 
perature transient approximates (for some seconds at 
least) tq that produced in a solid having, constant 
thermal coefficients, bounded internally by a circular 
cylinder and unbounded externally, initially all at a 
uniform temperature, to which a pulse of heat is sup¬ 
plied instantaneously and uniformly at the ijiitial 
time all over the inner surface, that surface being 


CONFIDENTIAL 



















































106 


HEATING OF GUNS DURING FIRING 


thereafter insulated. The transient is given by 


6 = 


2 H 

7 rpC 


/: 


Ji(ua)Y 0 (uT ) — Yi(uo)J 0 (ut) 
Jfiua) + Yfiud) 


( 10 ) 


du 


in which 0 is temperature rise, H is heat input per 
unit area, pc is volume specific heat, a (= k/pc) 
is diffusivity, t is time, a is mean bore radius, 
r is radial distance, and the J and Y are Bessel 
functions. 542 

A thermocouple assembly 106 designed at the Geo¬ 
physical Laboratory fitted into a hole of small dia¬ 
meter drilled radially to within a fraction of an inch 
of the bore of a gun barrel as shown in Figure 2. The 
junction was formed at the bottom of the hole, and 
the electrical insulation Was made thin in order that 
the flow of heat should be altered as little as possible. 
By means of a Leeds and Northrup Speedomax or its 
equivalent, time-temperature tran sien ts were reco rded 
during the firing of single rounds in different guns. 
The junction was situated about % in. from the bore, 
and a maximum temperature rise ranging from a few 
degrees to 20 or 30 degrees centigrade was recorded 
after about 2 sec, depending on the gun, the location 
of the thermocouple, and the ammunition. 

In some cases, although not in all, the recorded 
curve was matched quite accurately by equation (10) 
or a similar equation which took into account the 
finite outside diameter of the barrel. Since the volume 
specific heat pc (which occurs in the diffusivity) can 
be considered known, the equation contains two 
parameters, H and the thermal conductivity k. Change 
of H corresponds to a change of temperature scale 
and. change of k to a change of time scale. When the 
experimental curve is matched by a suitable choice of 
the two scales, the values of the heat input H and the 
conductivity k are determined. Of course, if k is 
known in advance for the particular steel, the de¬ 
termination of H is more precise. 

The method of the embedded thermocouple is the 
simplest for measuring heat input just forward of the 
origin of bore. However a completely reliable design 
of embedded thermocouple is still to be sought. In one 
developed at Purdue University 321 for the Army Ord¬ 
nance Department the end of the thermocouple junc¬ 
tion was at the bore surface, but the thermoelectric 
voltage produced corresponded to the temperature 
over a zone of metal several thousandths of an inch 
thick, which is the zone where the temperature gra¬ 
dient is steepest. 


5 3 7 Heat Input by Strain Measurement 

During the course of the firings of caliber .50 air¬ 
craft barrels at Carderock, described in Section 7.5, 
thermal strain was observed in the oscillograph rec¬ 
ords. These showed that after the effect of the gas 
pressure had subsided, the strain settled down about 
15 msec after firing to a value which remained con¬ 
stant for at least 35 msec. This constant strain was 
considered to be a thermal effect and was used to 
determine heat input. The method was based on a 
theorem proved in Section 7.2.3, to the effect that 
circumferential strain at the outer surface of a hollow 
cylinder depends only on the quantity of heat in the 
cylinder and is independent of the (cylindrically 
symmetrical) distribution of the heat. 

It was assumed that this theorem could be applied 
to the aircraft barrels; final temperature rise was then 
computed from the thermal strain and the heat input 
was thus found. 106 The coefficient of linear thermal 
expansion of the steel was taken as 11.2X10 _6 /°C, 
the volume specific heat as 0.86 cal/cm 3 /°C. Single 
rounds in a nitrided and chromium-plated barrel gave 
results shown in Table 4. Since values are correct 


Table 4. Heat input by strain measurement, nitrided 
and chromium-plated caliber .50 aircraft machine-gun 
barrel fired single shot. 


Distance from 

Heat input 

breech (in.) 

(cal/cm 2 ) 

4.4 

7.4 ± 0.5 

10.0 

7.6 ± 0.2 

14.0 

6.5 ± 0.4 

24.0 

6.5 ± 0.4 


within 10 per cent, except near the origin of bore 
where the underlying assumptions of the method are 
not met, the results were very satisfactory. They 
could be further improved by testing for thermal 
strain alone, in which case higher amplification could 
be used. 

5 3 8 Other Methods of Measurement 

Of other possible methods of measurement of input 
or temperature, one that has been of interest recently 
is measurement by means of temperature-indicating 
coatings. 412 - 432 ’ 433,529,530 ' 531 Such paints are now avail¬ 
able covering a range of 80 to 800 degrees centigrade, 
well suited to external barrel temperatures. Single, 
double, and multiple-change colors have been devel- 


CONFIDENTIAL 









BORE SURFACE TEMPERATURES 


107 


oped. Those of the first kind change color irreversibly 
at a fixed temperature; for the double-change type, 
the first change is a reversible one taking place at a 
relatively low temperature; continued heating causes 
a permanent change at a much higher temperature. 
The multicolored paints show varying irreversible 
changes at different levels; the regions of color are, 
however, not so sharply defined. It is claimed that 
with careful handling, indicated temperatures will be 
within approximately 5 per cent of the actual ones. 

The paints are dependent for effectiveness on time 
of exposure to heat. Some have been standardized on 
a basis of 30 min heating, others on 10 min. At tem¬ 
peratures which are high for the particular paint be¬ 
ing used, however, the color changes may result from 
much shorter times of exposure. 433 In the case cf 
Thermindex paint E6 a transition which requires 10 
min at 450 C will occur in 4 to 5 sec at 680 C. For 
this paint it has been shown that, on a graph of tem¬ 
perature versus logarithm of heating time required 
for a color change, a simple linear relation holds over 
the time range investigated for the two lower temper¬ 
ature transitions, but that the relation for the upper 
temperature transition is more complex. 

5 3 9 The Thermal Analyzer 

Various aids to the study of temperature distribu¬ 
tions have been devised; in particular, a thermal 
analyzer has been developed at the University of 
California under the auspices of the OSRD Engineer¬ 
ing and Transition Office. A network of electrical 
resistors and capacitors is used to simulate the ther¬ 
mal system. In a typical unit eight radial positions 
and eight axial positions are provided; impulses are 
introduced at any one of eight points to represent heat 
input. The corresponding first eight resistors repre¬ 
sent the resistance of the laminar layer next to the 
surface. Temperatures can be taken at any of 64 net¬ 
work positions at which condensers are attached. All 
condensers and resistors are variable, making the unit 
applicable to different guns and firing conditions. 


e In the fall of 1943, Section A of Division A, NDRC held a 
symposium to review the current knowledge about the heat 
input to the bore surface of guns and the temperatures result¬ 
ing. The ten papers presented, together with useful contributed 
comment, were later edited and issued as NDRC Report 
A-201. 32 Some of this material was later expanded and issued 
in separate reports which are referred to elsewhere in this 
chapter. 

f In general the mathematical notation follows that of the 
original sources, and therefore is not consistent throughout 


54 BORE SURFACE TEMPERATURES 6 


5,4,1 Theoretical Methods of Determination f 

The temperature of the bore surface may be cal¬ 
culated from the heat input and the known laws of 
heat flow. In this section are summarized some of the 
methods and results that various workers have ob¬ 
tained. All are based on the same fundamental as¬ 
sumptions : (1) the principal source of heat transfer is 
forced convection from the gas, as described in Section 
5.2.2; and (2) the heat conduction problem is that of 
flow through a semi-infinite slab. For this flow we 
have the fundamental equation 


dT = _k_ d 2 T 
dt C p 3 dz 2 


( 11 ) 


where T is the temperature of the gun metal at depth 
z from the surface and time t, and A*, c, and p 3 are 
respectively the thermal conductivity, specific heat, 
and density of the metal of the bore wall. The ratio 
(k/cpg), which is called the 1 Thermal diffusivity,” is 
considered constant in most applications. To obtain 
the temperature of the surface T a , equation (11) is 
integrated under the boundary conditions 


At « = 0, = - A( Tg - T a ) (12a) 

At t = 0, T = T s = T ao (12b) 

where h is the heat transfer coefficient discussed in 
Section 5.2.2. This integration leads to 


Ts{t) - T S0 n 

= (7 TPsCk )- 1 ' 2 J 0 h(T 0 - T s ){t - ?n)~ 1,2 dm (13) 

where m is a variable of integration. This equation 
shows the characteristic difficulty of the calculation 
of bore-surface temperature: the heat input as defined 
in equation (2) is itself dependent on the desired tem¬ 
perature T s . Hence numerical or other approximate 
methods of solution must be sought. 

The five methods now to be described differ prin¬ 
cipally in the manner of calculating h, in the values 
chosen for the thermal constants, and in the methods 
of approximating the solution of equations (11) and 

(13). 


this chapter. In order to simplify the reading of individual 
sections, however, some changes have been made. In particu¬ 
lar, throughout this section a is used for diffusivity and T for 
temperature, although in some of the sources other symbols 
had been used. 


CONFIDENTIAL 





108 


HEATING OF GUNS DURING FIRING 


Method of Fulcher 32 

In this method of computation the heat transfer 
coefficient used was that given by equation (6), the 
simplified theoretical expression for smooth pipes. 
Then the rate of heat input Q (in calories per square 
centimeter) was obtained from equation (14), in 
which U' is the velocity in thousands of feet per 
second. 

Q = 8.9 ( P Uy \T 0 - T a ). (14) 

Since Q varies regularly with time, equation (13) was 
approximated by 

T s (m) - T sq 

= 3 4 ( 1 + (^) 

in which n is the number of small intervals into which 
the graph of Q as a function of time is arbitrarily 
divided and / is a function of the curvature of that 
graph. 

The method of calculation consists in determining 
p, U', and T g from experimental records of breech 
pressure and projectile position as functions of time 
during firing, by using the fundamental gas laws. 
From the known data graphs are set up for total 
energy, energy remaining in the powder gas, gas tem¬ 
peratures, and velocity of the powder gas, as partially 
illustrated in Figure 3. Then the values for p, U', and 
T 0 are used with equations (14) and (15) by the meth- 



Figure 3. The total kinetic energy which the powder 
gas would have had if it had lost no energy, versus time 
(upper curve). This curve is determined by summing the 
energies as indicated on the three lower curves. (This 
figure has appeared as Figure 3 in NDRC Report No. 
A-201.) 


od of successive approximations to obtain values 
for the rate of heat input Q and the surface tempera¬ 
tures T s at any time and for various positions along 
the bore. 

An important advantage of this method of calcula¬ 
tion is that it is readily adaptable to different experi¬ 
mental conditions. Thus it is not limited to any partic¬ 
ular assumption about variation of gas density or 
gas temperature, and it is possible to take preheating 
(as by engraving) into account. 

Method of Hirschfelder, Kershner, 
and Curtiss 32 

In this method the smooth pipe approximation for 
h was used, in the form of equation (5), with the exact 
theoretical expression for h. The ballistic variables 
p, £/', and T 0 were obtained by the detailed theoretical 
methods of calculation described in Chapter 3. A con¬ 
stant value of 3.4 was adopted for the term 2(irp a ck)~ 112 
and equation (13) was approximated by 

T a (t m ) — To 

m _ _ 

= 3.4E<?(fe-i)(v / C - ti -1 - \A» - tj). (16) 

That is, the time was divided into m small intervals, 
and the rate of heat transfer assumed constant through 
each interval. A step-by-step computation was car¬ 
ried out, using the value of Qitj-J calculated on the 
basis of the T s at the beginning of the interval. 

The more general case for temperature T x at a dis¬ 
tance x from the bore surface was solved on the basis 
of an analogous equation. The resulting step-by-step 
summation method of solution is indicated by equa¬ 
tion (17) 

m 

T x {t m )=T x { 0) +E<5r a (< - t H ) ■ 

[*(*£) - 4%f)]■ 1171 

in which a is diffusivity and the function E(y) is the 
error integral defined by 

2 f y 

E(y) = e~ ui du . (18) 

Existing data were similarly fitted to empirical 
relations to give equation (19) for total input to entire 
barrel up to time of ejection (hi) 

c , l/ro „ nd m 


CONFIDENTIAL 

























BORE SURFACE TEMPERATURES 


109 


and equation (20) for average heating per square inch 
of bore surface per round. 

K L/P[.09(r 0 -300) 4 / 3 ] AD 2/3 . 

4 X 27.6 8 (§=) ' 

In these equations T 0 is flame temperature on the 
absolute temperature scale, L projectile travel in 
inches, C the charge in pounds, D caliber in inches, A 
loading density in g/cm 3 , X m is equivalent length of 
chamber in inches and X m /X 0 the ratio of gun volume 
to chamber volume. Equation (20) makes clear the 
well-known relation that for guns and loading such 
that L/D, X m /Xo, and A are the same, the heating per 
unit area per round varies as the two-thirds power of 
the caliber. 

In connection with these methods of calculation it 
is of interest to note approximate results for near-sur¬ 
face temperatures that have been found by the class¬ 
ical theory of heat flow in a semi-infinite solid. Heat 
input values suitable for normal guns and machine 
guns and thermal constants generally accepted for 
gun steel were used. The temperature rise at distances 
near enough to the bore surface to make x/4at < 0.01 
are shown in Figure 4. The maximum temperatures 
attained by points at various distances from the sur¬ 
face are shown in Figure 5, together with the elapsed 
time to the maximum. The functions used in these 
calculations are, of course, discontinuous at the bore 
surface itself. 

Nordheim’s Method 48 

This method uses the preferred expression, equa¬ 
tion (4), for heat transfer in a rough tube, in which 
the constant X is fitted empirically to the experimental 
results for heat transfer. The ballistic parameters are 
derived theoretically from the simplified ballistic sys¬ 
tem described in Section 3.2.7. For this system, all 
velocity-time curves are the same if described in 
terms of a reduced-time variable r. It then becomes 
possible to solve equation (13) in terms of this reduced¬ 
time variable and a corresponding reduced-depth 
variable so that the results are applicable to guns of 
varying sizes and conditions of firing. The simplifica¬ 
tions in the ballistic system cause relatively small 
changes in the calculated heat transfers and surface 
temperatures. 

Fundamental experiments on erosion vents 27 ’ 260 - 479 
also formed part of the basis for the analysis. A knowl¬ 
edge of an experimental value of the heat input to at 




Figure 4. Temperature rise near the bore surface after 
firing of one round. 


least one gun is required. From such a measurement 
values of the friction factor X in guns of different cali¬ 
bers were derived. Because the measurements 71 of 



DISTANCE FROM BORE SURFACE IN CENTIMETERS 

Figure 5. Maximum temperature attained from firing 
of one round. 


CONFIDENTIAL 








































































110 


HEATING OF GUNS DURING FIRING 


heat input were still in progress when this method of 
calculation was developed, two values of the heat in¬ 
put were introduced, so that interpolation might be 
made later on the basis of more accurate experimental 
determinations. 

In the computations it was further assumed that 
the density of the propellant gases is constant 
throughout the whole space behind the projectile. 
Constant values of k, c, and p s were chosen, so that 
the value of the ratio 2/( 7 rp s ck) 112 was 3.16. 

The simplified or reduced ballistic system permit¬ 
ted the development of formulas and tables inde¬ 
pendent of the absolute size of the gun, relatively in¬ 
sensitive to variations in loading density lying within 
the normal range from 0.5 to 0.7 g per cc, and to other 
ballistic variables. The results 48 depend essentially on 
two parameters. The first is the “burning parameter” 
E which fixes the position of the projectile when the 
powder is all burnt. E may be calculated from the 
powder and gun constants, and then the position y b at 
all-burnt is given by equation ( 21 ), 


~ = [ 1/(1 - E)]V«-V ( 21 ) 

2/0 

in- which y is a reduced-position variable, the sub¬ 
script 0 indicating the initial position of the projectile. 
In preparation of the tables , 48 a value of 1.36 was used 
for 7 , although as pointed out in Section 3 . 3.2 a some¬ 
what lower value would be preferable. 

The second fundamental parameter is the “heating 
parameter” L. This variable is independent of time, 
contains the friction factor X, and may otherwise be 
calculated from the gun and powder constants by 
equation ( 22 ), where the symbols have their usual 
ballistic significance (Chapter 3 ). 


L = 


43.1Xc p C 3/4 ~| [" 1 

.Vkcp s (m 1,4 A l,2 F)J L 


— Appp l 

A 0 J 



( 22 ) 


There are three periods of time which have special 
significance in temperature studies—prior to all-burnt, 
after the burning of the powder but preceding ejection 
of the projectile, and after the ejection. For posi¬ 
tions back of y b , the heat input is greatest during 
the first period, smaller in the second, and relatively 
slight after ejection. For positions toward the muzzle 
from y h , heating begins in the second period and is 
important in the third. It has been remarked that 
uncertainties as to the behavior of the gas stream 
following ejection make the figures given in NDRC 
report A-262 less reliable; they have, however, com¬ 


pared reasonably well with experimental results, as 
brought out at the end of Section 5 . 2 . 2 . 

On the reduced basis the fundamental heat-transfer 
equations ( 11 ) and ( 12 ) become equations (23) and 
(24), 


dT 

dr 

dT 

ar 


1 d 2 T 

2 ar 2, 

-H(Tg-T') 


(23) 

(24) 


The reduced heat-transfer coefficient H (r) is the prod¬ 
uct of three factors, as shown in equation (25). 


H(r) = 



(25) 


These factors are, respectively, the position in the 
gun X s /X 0 , the time-independent heating parameter 
L, and the general function /(r), which depends only 
on the burning parameter E. Thus a general solution 
of the problem has been obtained. 

The heat-transfer function /(r) plotted against r 
gives an asymmetric bell-shaped curve. A simpli¬ 
fied method of finding approximate maximum tem¬ 
peratures near the beginning of travel involves use 
of a straight line to replace this curve between the 
origin and the point corresponding to “all-burnt.” In 
most cases results obtained by this method will be 
within 30 centigrade degrees of those given by the full 
tabular method, usually considerably closer. 

Tables have been issued 48 for the temperature rise 
corresponding to various values of L, assuming flame 
temperatures of 2500 K for “normal” guns and 2795 K 
for machine guns. A slightly different treatment was 
used for machine guns because their ballistics are 
somewhat different from those of larger guns. These 
tables contain values corresponding to different 
depths below the bore surface. Separate tables are 
given for times prior to all-burnt and after burning 
has been completed. 

The main tables are based on the assumption that 
the unburned powder grains are distributed uniformly 
in the space behind the projectile. An auxiliary table 
can be used to obtain values corresponding to the 
assumption that all the unburned powder remains in 
the chamber. 

Correction functions are presented for changes in 
the flame temperature of the powder and the initial 
temperature of the bore surface. It has been pointed 
out , 113 however, that it is undesirable to extrapolate 
beyond a flame temperature of 3000 K for normal 
guns. In several cases results for hotter powders that 


CONFIDENTIAL 





BORE SURFACE TEMPERATURES 


111 


have been compared with those obtained by numeri¬ 
cal integration have shown larger variations (as much 
as 100 degrees than for cooler powders. 


Method of Hicks, Thornhill, and Ware 

In developing this method the heat transfer coeffi¬ 
cients were assumed known from the smooth-pipe 
expression and empirical values of the ballistic quan¬ 
tities. By application of the boundary conditions 
given by equations (26), 

t = 0, T = T 0 ) 

z = 0, dT/dz = - (h/k)(T g - T) f (26) 
z —> oo T 7 —>0 J 

a convenient numerical solution 348 of the heat conduc¬ 
tion equations (11) and (12) was obtained. As orig¬ 
inally described it may be outlined as follows: 

Dividing up the time interval over which heat 
transfer takes place into a number of equal small in¬ 
tervals 8t , equation (11) may be replaced approxi¬ 
mately during one of these intervals by equation (27), 


Ti(Z) - Tp(Z ) 

8t 

= (1/2) [To(Z) + UZ)], (27) 

where the T s are the temperature-space distributions 
at the beginning and end of the interval. By applying 
the definition of q given by equation (28) 


= 2 cp 8 
q k8t ’ 

equation (27) reduces to equation (29). 

-~(T 0 + TO = q\T 0 + TO - 2a*3V 


(28) 

(29) 


After proper consideration of the boundary condi¬ 
tions, the temperature distribution at the beginning 
of the (n + 2)nd interval 8t may be written 


T 0 (Z) = e~« z [ao + a x {2qz) 

+ ■■■+ -£(2 9 r)J. (30) 

The temperature at the end of this interval is given 
by a similar relation with coefficients b 0 to b n+v 
Coefficient bp may be expressed in the form 


hpTgp T~ hiTgi _ hp -f- kq 

hi -j- kq hi -j- kq 

+ fry [«o + «i + * * • a«] (31) 


The other 6’s are functions of the as alone. Starting 
from the initial temperature T 0 (Z) = 0, these recur¬ 
rence relations enable one to calculate the tempera¬ 
ture distribution at the end of any subsequent interval. 

If instead of using the varying values of h(t ) and 
T a (t) and integrating by a step-by-step method we 
assume that these quantities may be replaced by 
some form of average values h and T g , the heat con¬ 
duction equation may be solved analytically, as 
shown by Thornhill and Ware 398,403 and by Hobstet- 
ter. 124 The latter has written the result in the form 
T - T„ „ J Z \ . £z + £«< 


[‘- E XwS +¥''«')} 


(32) 


In this equation a is the diffusivity and h the average 
heat transfer coefficient. The abbreviation Erf stands 
for the error function; 521 for example, 


Erf 


2 \^ai 


=- 2 = r 

V * J o 


Z/2\/ at 

e~ u2 du. 


(33) 


T g is likewise taken as an average; the other symbols 
are as previously introduced. 

In applying this method, Thornhill and Ware have 
considered time to projectile ejection as time of trans¬ 
fer of heat, the average T g to be the mean of the 
initial and final gas temperatures, and the average h 
to be half the maximum numerical value of h(t). As 
compared with their more accurate method outlined 
above, they have found the values for heat transfer to 
be correct within a few per cent, and the maximum 
surface temperatures to be about 15 per cent too low. 
Since the calculations were made for the origin of 
rifling, the input after ejection is a minor factor; how¬ 
ever for points farther down the barrel, ignoring this 
would lead to a considerable error. 


Hobstetter’s Method 124 

In an effort to obtain a truer picture of the situa¬ 
tion by a different choice of parameters, the analytic 
expression of the previous method was combined 
with the heat transfer coefficient and other ballistic 
quantities as calculated by Nordheim’s simplified 
system described above. For points near the origin of 
rifling, and under the assumption of equal distri¬ 
bution of unburnt powder, the gas temperature de¬ 
creases slowly and therefore it was concluded that a 
good approximation is given by taking this equal to 
the flame temperature of the powder. 


CONFIDENTIAL 










112 


HEATING OF GUNS DURING FIRING 


For the duration of heating the time to maximum 
temperature was expressed graphically in terms of 
the burning parameter E. Finally, for the heat trans¬ 
fer coefficient equation (34) was adopted. In it h is 
the transfer coefficient used by Nordheim, 

h' = (34) 

a is the ratio r/t between Nordheim’s reduced time 
and time in the ordinary scale, k, c, and p s have the 
same meaning as in equation ( 11 ). 

This method was applied to studies of caliber .50 
and 37-mm guns. The former computations were 
carried out in conjunction withmetallographic studies 
of bore surface erosion, which are discussed in Chap¬ 
ter 13. 


Comparison of the Methods 

A comparison of the results by the different meth¬ 
ods is desirable; it has been difficult to find guns for 
which calculations have been made by all the methods. 
For two guns the values given in Table 5 have been 


Table 5. Bore surface temperatures calculated by vari¬ 
ous methods. 




Method 



Simplified 

Nordheim 

Hob- 

Hirschfelder 

Gun 

Nordheim 

tabular 

stetter Kershner et. al. 

37-mm M3 

570 C 

590 C 

532 C 

555 C 

8-in. Mk VI 

1017 C 

1022 C 

1014 C 

1210 C 


calculated. It maybe remarked that in general, results 
are not so consistent as the 37-mm gun temperatures 
would imply; on the other hand, the Hirschfelder re¬ 
sults are usually in better agreement than the 8 -in. 
gun results would indicate. 

Temperature in Continued Fire 

The preceding section on calculation of bore sur¬ 
face temperature dealt with actual instantaneous 
temperature, and had in view chiefly a single round. 
In the firing of a continuous burst a concept of im¬ 
portance is average temperature. It is best defined as 
the temperature that would be produced if the heat of 
each round entered in a uniform way over the whole 
time between that round and the next, instead of al¬ 
most at once after the instant of firing. The average 
temperature rises smoothly and steadily as the burst 
goes on whereas the actual temperature at or near the 


bore surface rises far above the average almost im¬ 
mediately after a round is fired and falls below the 
average during most of the time between rounds. At 
a distance from the bore surface the actual and aver¬ 
age temperatures are practically the same in rapid 
fire . 106 The progress of heating of the whole barrel is 
described at some length in Section 5 . 5 . 

The average temperature in the neighborhood of 
the bore surface tells the general temperature level, 
as distinguished from the momentary sharp peak of 
temperature after each round. The latter is important 
chiefly for chemical effects and bore surface melting, 
the former rather for deeper mechanical effects fol¬ 
lowing upon bullet impact, as described in Section 
13.1.2. 

Average bore surface temperature can be deter¬ 
mined from measurements of temperature during a 
burst taken at several distances from the bore. Such 
measurements were made in caliber .50 barrels at the 
Geophysical Laboratory by means of the embedded 
thermocouples described in Section 5.3.5. The read¬ 
ings were taken by photographing millivoltmeters and 
a clock, or by pen and ink recorders. In one case, four 
thermocouples were inserted at the origin of bore, at 
distances from the bore ranging from J/f 6 in. to the 
outer surface. Soon after the beginning of the burst, 
temperatures were rising at nearly the same rate at 
all four thermocouples (see Figure 8 , couples 4.35 in. 
from breech). The heat input H per unit time, through 
a bore surface area of unit length axially and extend¬ 
ing one radian circumferentially, could then be cal¬ 
culated by equation (35), 


II 


e(ri) - 0(r 2 ) + 



(35) 


in which 0 is temperature, n and r 2 are the distances 
from the axis of two thermocouples, k is conductivity, 
a is bore radius, pc is volume specific heat, and t is time. 
The average bore surface temperature was given by 

0(ri) + H L Ta~ fail ? c s§ dsd '- ( 36 > 

In these two formulas the effect of any possible 
flow of heat in the axial direction has been neglected. 
If axial flow is not neglected, additional terms appear 
in both formulas. An estimate of the axial flow can be 
made from the readings of three thermocouples . 106 

The principal term in the numerator of equation (35) 


CONFIDENTIAL 










BORE SURFACE TEMPERATURES 


113 


in the calculation of the rate of heat input H is ©(r^) 
— 0(r 2 ),the temperature difference at a given moment 
between thermocouples at two different distances 
from the bore. In the example cited above the differ¬ 
ence amounted to about 80 degrees centigrade be¬ 
tween 14 in* and Y% in. from the bore. It is doubtful, 
however, whether the embedded thermocouples were 
capable of measuring such a difference very accu¬ 
rately. 

The amount by which the average bore surface 
temperature exceeds the actual temperature just be¬ 
fore a new round may be estimated by 1.46# X 
(n/ irpck) ^, in which B is the heat input of the new 
round per unit area and n is the rate of fire of the 
burst. 106 

It is possible .to calculate average temperatures in a 
burst, and for that matter actual temperatures not too 
close to the bore, by in effect accumulating the tem¬ 
peratures of the separate rounds as expressed in equa¬ 
tion (10) or analogous equations, provided the section 
in question is not too close to the origin or the muzzle, 
and the outer surface of the barrel is nearly cylindri¬ 
cal in the neighborhood of the section. Besides the 
dimensions of the barrel and the thermal coefficients 
of the steel, it is necessary only to have some knowl- 



Figure 6. Calculated and experimental temperatures 
in 3-in./50-cal. Mark 22 Naval gun firing 90 rounds at 
the rate of 6 rounds per minute. Allowance is not made 
in the calculation for heat loss through the outer surface. 
(This figure has appeared as Figure 70 in NDRC Report 
No. A-434.) 


edge of the heat input per round. This can be gained 
from measurement at the outer surface during a burst; 
or from measurement of the input of a single round 
coupled with the assumption that heat input de¬ 
creases in proportion as average bore temperature 
approaches a fixed “limit temperature” of about 
1000 C; or finally, without measurement, from Nord- 
heinTs tables or an equivalent theoretical calculation 
(see Section 5.4.1). 

Figure 6 shows calculated average temperatures in 
a 3-in. gun firing 90 rounds at a rate of 6 rounds per 
minute. These temperatures are drawn in full line, for 
the bore surface and the outer surface, at a position 
near the muzzle and another about midway along 
the barrel. Measured outer surface temperatures at 
the same positions are drawn in dotted lines. The cal¬ 
culations were based on measurements of single round 
heat input, and the assumption of a limit temperature 
of 1000 degrees near the middle of the barrel and 700 
degrees near the muzzle. It was assumed also that no 
heat escaped from the outer surface. If an allowance 
had been made according to known rules for the es¬ 
cape of heat there, the calculated and measured outer 
surface temperatures would have come very near to 
agreement. 106 

5 4 3 Experimental Methods of 

Determination 

Bore Surface Thermocouple 

A means of recording the transient temperature at 
the bore surface of a gun was devised by a German 
investigator. 492 The measurement was made with a 
thermocouple assembly that could be inserted in a gun 
wall so that the thermo-element formed part of the 
gun bore surface. This thermo-element consisted of 
an oxidized nickel wire that passed down a fine hole in 
a steel plug and a thin layer of nickel that had been 
plated on the exposed end. This layer of nickel bridged 
the layer of nickel oxide that insulated the wire from 
the plug, in which it was a tight fit. There is not 
available any record of the extent to which this 
method was used. 

Fusion Temperature of Erosion Products 

Many of the chemically altered erosion products 
that occur on the bore surface give evidence of having 
been liquefied, as is brought out in Section 12.6. 
Hence a determination of the temperature of incip- 


CONFIDENTIAL 












114 


HEATING OF GUNS DURING FIRING 


ient fusion of such materials gives a lower limit to the 
temperature of the bore surface. An experimental 
method for such a determination has been worked out 
in a preliminary way. 105 

A few milligrams of the bore surface deposit were 
sealed in a very small, thin-walled, well-evacuated 
silica-glass container whose volume was about 20 
mm 3 . The container was heated in a vertical electric 
furnace (2.5 cm bore), which had been preheated to 
the temperature at which it was desired to heat the 
charge. 

The furnace temperature was measured with a 
platinum-platinrhodium thermocouple junction lo¬ 
cated a few millimeters above the bottom of a 7-mm 
silica glass tube, closed at its lower end, which was 
placed at the hottest part of the furnace. 

The sample in its evacuated container was lowered 
on a platinum wire to the bottom of the silica tube, 
left there for a certain number of seconds (long enough 
to attain a temperature which was certainly not more 
than 5 or 10 degrees lower than the furnace tempera¬ 
ture), and was then lifted out and the contents ex¬ 
amined for signs of melting. If no melting had taken 
place, a new sample was used for the next trial at a 
higher temperature. Suitable criteria for melting had 
to be found for each substance studied. X-ray exam¬ 
inations were made of the phases in the material be¬ 
fore and after melting. 

Interpretation of the results is difficult because of 
the limited amount of experimental data. A residue 
from decoppering (see Section 11.4.3) showed a fusion 
temperature of 1160 ± 40 C. As now set up, the meth¬ 
od unfortunately cannot be used for temperatures 
above 1250 C. because of softening of the silica glass 
container. 

544 Applications to Conditions of Firing 

Effect of Preheating 

As stated in Section 5.2.2, the methods of calcula¬ 
tion just discussed have ignored heat input from 
sources other than transfer from the hot gases. How¬ 
ever, an indication of the effects from engraving and 
frictional forces has been given there and in Section 
5.3.8. 

The reality of a high temperature at the interface 
between a projectile and the gun bore has been dem¬ 
onstrated in the case of caliber .50 bullets. 72 By using 
the gilding metal of the bullet jacket in contact with 
the steel of the bore surface as a sliding thermocouple, 


measurements of the temperature of the interface 
were made while the bullet was in motion in a short¬ 
ened caliber .50 barrel. A temperature of 860 C was 
recorded for the apparent average temperature over 
the area of contact between the bullet and the bore; 
and very likely some spot reached a higher tempera¬ 
ture. 

The question naturally arises as to the effect of 
such preheating® on the bore surface temperature. 
When the time of preheating is very small (less than 
half the burning time) as in the case cf preheating 
resulting from either engraving or bore friction, the 
increase of maximum temperature due to gas heat¬ 
ing is relatively low. 48 For a 37-mm gun the increase 
per preheating unit of 1 cal/cm 2 for the neighborhood 
of the origin of rifling is 25 degrees, centigrade. As 
mentioned at the end of Section 5.3.8, one set of as¬ 
sumptions led to an estimate of approximately two 
such units for the maximum preheating input. There¬ 
fore, we may be safe in saying that for this gun the in¬ 
crease of maximum bore surface temperature from 
preheating due to the rotating band does not exceed 
50 degrees centigrade and it is probably considerably 
less. 

Effect of Bore Surface Material 

When a liner is inserted in a gun, or when the bore 
surface is plated, heat input and resulting bore sur¬ 
face temperature are correspondingly affected. The 
metals now of most interest as liner materials are 
chromium-base alloys (Chapter 17), molybdenum 
(Chapter 18), and Stellite No. 21 (Chapter 19). The 
one commonly used in plating is chromium (Chapter 
20 ). 

In plated guns it is essential to have the depth of 
plating sufficient to prevent thermal alteration of the 
underlying steel (see Section 13.3 and Section 31.5). 
When this is the case it is necessary in calculating 
heat input to the bore surface from a single round to 
consider only the thermal constants of the plated metal 
and its surface roughness as compared with gun 
steel. After a few rounds have been fired in a gun 
there appears to be little difference between metals 
in regard to the latter factor. 

K It is important to note that the correction functions given 
in Table 38 of NDRC Report A-262 48 and applied in formula 
6.16 of that report applies to initial temperatures that are 
fairly uniform throughout a layer of the bore surface consider¬ 
ably thicker than the maximum depth considered in the 
tables. This preheating might be from a previous round or 
from an external source, such as the sun. 


CONFIDENTIAL 




RATES OF HEATING UNDER RAPID FIRE 


115 


Since liners are much thicker than plating, it is 
sufficient for purposes of comparison of temperatures 
at the bore surface to base calculations entirely on the 
thermal constants of the metals occurring there. Un¬ 
fortunately these are not too well established, even 
for gun steel. Also, since both conductivity and vol¬ 
ume specific heat are functions of temperature, there 
is always some question as to what values to adopt. 
Values for steel and stellite are given in Table 3 of 
Chapter 19. Values for chromium and molybdenum 
may be derived from data given in Chapters 17 and 
18. The melting temperature of molybdenum is so 
high, 2620 C, that liners made of this metal will not 
melt with any propellants now in service. 

If average values for the constants are assumed, 
calculations indicate that the bore surface tempera¬ 
tures reached in gun steel and in stellite-lined guns 
differ by very little. For the caliber .50 erosion testing 
gun (Section 11.2.1), calculated 124 values using 
FNH-M2 powder are 1670 C and 1680 C, respectively, 
for the temperature after a single shot. Surface melt¬ 
ing was observed in both cases. 76 - 77 With IMR powder 
the calculated results are 1310 C and 1325 C. Here it 
is important to note that the lower end of the fusion 
range of stellite is exceeded slightly while the melting 
point of gun steel is not reached; but yet surface melt¬ 
ing of the stellite was not observed after 500 rounds. 
Gun steel did not melt when one round of IMR 
powder was fired, but liquefaction of the steel which 
had been chemically altered by the powder gases 
occurred after a number of rounds. 124 It is not sur¬ 
prising that, in the case of continued fire with IMR 
powder, steel melts while stellite does not, for the 
melting range of altered gun steel is much lower than 
that of stellite, as is brought out in Section 12.5.3. 

The bore surface temperatures for a stellite-lined 
37-mm gun for three FNH-powders have also been 
similarly computed. The estimates are 980 C for Ml, 
1310 C for M5, and 1330 C for M2; the last powder 
is considerably hotter than M5, but a smaller charge 
is sufficient to give the desired velocity of 2,900 fps. 
These figures predict melting of the stellite when the 
hot powders are used; this has been verified experi¬ 
mentally in firings of a 37-mm stellite liner at 
Aberdeen, as described in Section 33.2. 

Temperatures in a nitrided, chromium-plated air¬ 
craft barrel, or in one containing a 9-in. stellite liner, 
are not remarkably different from those in a steel 
barrel, especially at the higher levels. Chromium 
plate permits more heat to enter in a single round or a 
short burst, but in longer-continued fire it has a con¬ 


trary effect. Average temperature at the bore of a 
stellite liner, at the origin, is at first considerably 
higher than in steel, but by the 200th round of a long 
burst, the difference is in great part erased. 106 

Effect of Obturation 

The “blow-torch” effect of gas leakage resulting 
from imperfect obturation is particularly likely to oc¬ 
cur in worn guns. For banded projectiles, the danger 
of gas washing is greatest at the beginning of the 
travel of the projectile, before the rotating band has 
been engraved. 48 Later the band forms a more effec¬ 
tive seal, and the increased velocity of the projectile 
reduces the time for gas washing. There are two chief 
reasons for the relatively large effect of the escaping 
gases; first, that the velocity of the gas stream is very 
high—equal to the velocity of sound—and second, 
that the friction is large in narrow and irregular slits. 
The result of gas washing is localized melting leading 
to scoring of a gun, as discussed in Section 13.4.1. 

It is clear that in the case of pre-engraved projec¬ 
tiles the possibility of gas leakage is always present, 
and that this would be accentuated by any imperfec¬ 
tions in their manufacture. Experiments with caliber 
.50 projectiles, described in Section 31.4.2, indicated 
an increase in erosion with increased leakage area. 
Whether the same conclusion holds for larger projec¬ 
tiles is not yet known. The only experimental evidence 
is that obtained with the 37-mm gun, T47 (Section 
31.7). Erosion was disappointingly rapid; but it is not 
certain that gas leakage was an important factor. 

55 RATES OF HEATING UNDER 
RAPID FIRE 

551 Machine Guns 

The heating of the caliber .50 machine gun barrel 
has been studied far more than that of any other gun. 
In part, this is because the caliber .50 was used as a 
laboratory test piece through which the cause and 
cure of erosion in other guns might be understood, 
and in part for its own importance as an aircraft 
weapon. The study included both the 28-lb, 45-in. 
heavy barrel and the 10-lb, 36-in. aircraft barrel; 
nitrided and chromium plated and stellite lined barrels 
as well as steel barrels; single rounds, bursts, and con¬ 
tinued fire under various firing schedules; different 
positions on the outer surface and inside the barrel 
wall; different ammunition; and different methods of 


CONFIDENTIAL 



116 


HEATING OF GUNS DURING FIRING 







Figure 7. Diagrammatic representation of the successive stages of heating of a cross section of a machine gun barrel 
during the firing of a single long burst. At first the high temperature is limited to a thin annulus at the bore surface The 
heat spreads out radially with a very steep gradient continuing close to the bore surface. If the barrel is allowed to cool for 
a short time (last diagram on the right), the steep gradient disappears and the temperature becomes almost uniform 
throughout the cross section. 


cooling. The work was done at Aberdeen Proving 
Ground, the Crane Company, the Geophysical Labo¬ 
ratory, Leeds and Northrup Company, the Naval 
Bureau of Aeronautics, North American Aviation 
Corporation, the University of California and at 
several places by the British. 11 

The process of heating, which is shown diagram- 
matically in Figure 7, is as follows: With each round 
that is fired through a caliber .50 barrel, heat enters 
through the bore surface. At the first round the input 
amounts to about 5 cal/cm 2 per round at the muzzle, 
increasing at first slowly and then nearer the breech 
more rapidly to about 10 cal/cm 2 per round at the 
origin. In the half inch or so of space between the ori¬ 
gin and the mouth of the cartridge case the input may 
considerably exceed 10 cal/cm 2 where the bore sur¬ 
face is inclined to the axis of the barrel, forming part 
of the surface of a cone instead of a cylinder. From 
the mouth of the cartridge case back to the breech, a 
distance of nearly 4 in., the walls of the chamber re¬ 
ceive little or no heat directly, for the cartridge case is 
ejected so quickly that not much heat can pass 
through its walls. (In fact, once the gun becomes hot, 
it may be that the net effect of each freshly loaded 
cold cartridge is to extract heat from the chamber 
walls. The input may then be positive in one direction 
from the cartridge case mouth and negative in the 
other.) The rounds after the first convey decreasing 
amounts of heat to the barrel; the input to a barrel at 
500 C is everywhere roughly half what it is to one at 
room temperature. 

It takes a matter of seconds for any considerable 
part of a pulse of heat received at the bore surface to 
make its appearance at the outer surface. On the 
other hand at least 6 or 7 rounds per second are fired 
in the heavy barrel, and as many as 20 rounds per 
second in the aircraft barrel in the M3 gun. The 
average temperature (Section 5.4.2) therefore builds 


h For references see the bibliography in Report A-434. 106 


up rapidly near the bore surface, and may exceed the 
temperature at the outer surface by hundreds of de¬ 
grees centigrade. During firing the greatest flow of 
heat is therefore radial, directly outward from the 
axis of the barrel, under the influence of these steep 
gradients. To a less extent flow parallel to the axis 
also takes place, especially from the region of the ori¬ 
gin back toward the thermally shielded chamber 
walls, and wherever the outer contour of the barrel 
changes abruptly. 

For a barrel firing in still air, the escape of heat 
through the outer surface is always small compared 
with the rate of entry through the bore, even when 
the outer surface is as hot as 800 C. If the barrel is 
cooled by a strong air blast, however, the rates may 
become nearly equal at such high temperatures. 

Within a few seconds after firing stops the radial 
temperature differences subside almost completely, 
especially in still air or where the barrel wall is thin. 
Axial differences on the other hand, the equalization 
of which would require flow through a long path of 
metal, may remain considerable for minutes. In partic¬ 
ular, in still air, the origin section may continue 250 
centigrade degrees cooler than the forward part of an 
aircraft barrel for 3 min after a long burst. 

The temperature levels attained at a given section 
depend chiefly on the heat input per round, the num¬ 
ber of rounds, and the wall thickness at the section. 
The aircraft barrel, as it happens, is so tapered that 
wall thickness is in fairly constant ratio to heat input 
along most of the length. From a few inches beyond 
the origin to a few inches short of the muzzle, there¬ 
fore, average temperatures after a given number of 
rounds are fairly uniform. They are lower toward the 
origin because of the flow to the chamber walls, and 
toward the muzzle because of the bulge in the last 3 
in. of the barrel’s length. In the heavy barrel, tempera¬ 
tures are much lower in the rear half of the length 
than in the corresponding part of the aircraft barrel 
because of the much greater wall thickness. In fact 


CONFIDENTIAL 







RATES OF HEATING UNDER RAPID FIRE 


117 



Figure 8. Temperature measurements at 10 different positions in a caliber .50 heavy machine gun barrel firing a burst 
of 135 rounds of AP ammunition, M2. (This figure has appeared as Figure 34 in NDRC Report No. A-434.) 


the bulges in that region were no doubt placed there 
largely as thermal sinks. 

Figure 8 shows measurements of temperatures at 
10 different positions in a caliber .50 heavy barrel 
firing a burst of 135 rounds of AP ammunition, M2. 
These were made with thermocouples attached at the 
surface and others embedded as described in Section 
5.3.5. The barrels were either solid steel or contained 
short steel liners. The figure is a composite of several 
firings at rates ranging from 400 to 500 rounds per 
minute, but the temperatures are plotted as functions 
of the round instead of the time, for a more equal 
comparison. One thermocouple was at the outer sur¬ 
face, 2.50 in. from the breech, over the chamber; four 
were at different distances from the bore at 4.35 in. 
from the breech (the origin of bore section); two were 
at 9.35 in. from the breech; and one was embedded at 


6.35, 25, and 44 in. from the breech. The figure 
illustrates, among other things, the low temperatures 
over the chamber; the manner in which the tempera¬ 
ture differentials at any one section are established in 
the first 40 or 50 rounds of a burst and are thereafter 
maintained rather constant; and the magnitude of 
these differentials. 

An example 106 of maximum temperatures at dif¬ 
ferent points of a steel aircraft barrel after firing five 
100-round bursts with a 2-min cooling interval be¬ 
tween bursts is given in Figure 9. Maximum tempera¬ 
tures (centigrade degrees) after five 100-round bursts, 
with 2-min cooling between bursts, from a barrel ini¬ 
tially at 20 C, are written on the figure at 16 places. 
The distances of these places from the breech are4.35 
in. (origin of rifling), 8.35 in., 24.00 in., and 35.00 in. 
(1 in. from muzzle). The distances from the bore are 


z 3 

T 8 



^540° 1 

1 1 - 1 - 

- 1 -1- 


W~ 


U-730° 

^ 7IO* 670° \ 



^690° 

£n... j 

^760° 

r~' 

^---* 

^730° 690°^ 

- - 

ORIGIN OF , 
rifling's^ 

\ ^ 70 - 

920 

900 

\ ^760° 710° ^ J 

820 760 

p 

_L 

! 

_1_1_ 1 

1 1 


0 4 

A. " 

8 

IR Q “ 

12 16 20 24 28 32 31 

5 36 


DISTANCE FROM BREECH IN INCHES 


Figure 9. Maximum temperatures at different points of a steel aircraft barrel after firing five 100-round bursts, with 2- 
min cooling between bursts. (This figure has appeared as Figure 57 in NDRC Report No. A-434.) 


CONFIDENTIAL 









































118 


HEATING OF GUNS DURING FIRING 


0.01 in., Jig in., Y in. and the outer surface distance. 
The temperatures *4 in- from the bore at 4.35 and 
8.35 in. from the breech are experimental and are 
770 C and 810 C, respectively. The others are esti¬ 
mated on the basis of other firings. The accuracy is 
±30 degrees centigrade. At the same four distances 
from the bore, temperatures are nowhere more than 
40 degrees higher than those shown. Maximum tem¬ 
perature much closer to the bore than 0.01 in. is not 
at present known accurately. 

Here may be seen the uniformity of temperature 
over the middle two-thirds of the barrel; the lower 
overall level at the origin and muzzle; and the de¬ 
crease forward from the origin of difference between 
average temperature at the outer surface and 0.01 in. 
from the bore owing to decrease in heat input per unit 
area and in wall thickness. These maximum tempera¬ 
tures were not reached simultaneously. At 0.01 in. 
from the bore the time interval by which they fol¬ 
lowed the passage of the last of the 500 bullets was 
about 0.008 sec, at Ye in. 0.1 sec, at Y% in. 0.4 sec, and 
at the outer surface from 2 to 5 sec, depending on the 
wall thickness. 

The very high value of 920 C for maximum tem¬ 
perature 0.01 in. from the bore at the origin is note¬ 
worthy. This is a calculated rather than an experi¬ 
mental value. The highest temperature actually meas¬ 
ured at the Geophysical Laboratory in a caliber .50 
steel barrel was the 810 C shown 4 in. forward of the 
origin at Ye in. from the bore. The highest tempera¬ 
ture measured in any caliber .50 barrel was 900 C, 
Ye in. from the bore at the origin of rifling of an air¬ 
craft barrel having a stellite liner, at the end of a con¬ 
tinuous burst of 440 rounds. 106 

Forced cooling is discussed in Section 5.7. It may 
be remarked here that the various methods (external 
liquid cooling excepted) are effective in reducing 
average barrel temperatures principally in intermit¬ 
tent fire. In a continuous burst it is not easy to take 
heat from the barrel and remove it to a distance or 
convert it to a less objectionable form of energy, as 
rapidly as it is supplied through the bore surface. 

Temperature in caliber .30 machine guns has been 
studied at Aberdeen Proving Ground, 175 and in a 
0.303 in. Bren gun by the British. 398 

Automatic Cannon 

The British reported 428 heating rates in a 20-mm 
Hispano Mark V gun, in connection with cook-off 
trials. Average rise in external chamber temperature 


for a given number of rounds, fired automatically at a 
rate of approximately 140 rpm, is shown in Table 6. 

Table 6. Heating of chamber of 20-mm Hispano gun, 
Mark V, fired long bursts at approximately 140 rpm. 


No. of rounds 175 200 225 250 275 300 

Temperature rise (C) 234 279 302 325 349 382 


These figures may be compared with temperatures 
recorded when firing 25-round bursts at 1-min inter¬ 
vals. The equivalent of 24 such bursts produced a 
peak temperature of 239 C; 15 led to a figure of 202 C. 

Tests by the Bureau of Ordnance with 20-mm AA 
guns, while not giving specific temperatures, indicated 
that 120 rounds of automatic fire result in a danger¬ 
ously high temperature. For 40-mm air-cooled guns, 
45 rounds fired in 8-round bursts with 15 sec pauses 
between bursts resulted in a temperature too high for 
safety. 

For 40-mm water-cooled guns, tests made at the 
Naval Proving Ground gave the results shown in 
Table 7 for increase in temperature of the coolant. A 


Table 7. Temperature of coolant of water-cooled 40-mm 
gun after having been fired bursts of different lengths. 


No. of rounds 

50 

80 

100 

150 

Time in seconds 

25 

45 

50 

80 

Temperature rise (C) 

25 

41 

55 

70 


change in gun elevation from 0 to 60 degrees reduced 
the firing rate slightly but had little effect on the 
water temperature. 

5 5 3 Medium Caliber Guns 

Aberdeen Proving Ground Tests 

The heating of guns of medium caliber has been 
studied for a number of years at Aberdeen Proving 
Ground. 1 In one investigation temperatures resulting 
from rapid fire in a 105-mm A A gun, Ml, were meas¬ 
ured. 178 Following 20 preliminary rounds, series of 58, 
66, 28, and 16 rounds were fired at varying rates. 
Temperatures were recorded by means of thermo¬ 
couples welded to the gun barrel at distances of 0.95, 
7.35, and 13.7 ft, respectively, from the muzzle. Maxi¬ 
mum values reached were 285 C, 300 C, and 280 C, 
respectively. The overall time for the 168 rapid-fire 
rounds was approximately 42 min. 


* These studies have been summarized and discussed in 
Ballistic Research Laboratory Report No. 104. 186 


CONFIDENTIAL 










RATES OF HEATING UNDER RAPID FIRE 


119 


From the heating effects, the temperature rise of 
the gun for rates of fire of 5, 10, and 15 rpm was com¬ 
puted and the results plotted. The times for the tem¬ 
peratures to reach 600 degrees centigrade above at¬ 
mospheric near the origin of rifling were found to be 
approximately 100, 35, and 25 min, respectively. 

Similar tests were made with a 3-in. A A gun 
M3. 179 Two thermocouples were welded to the gun, at 
1.08 and 5.21 ft from the muzzle. After 129 rounds had 
been fired in about 8 min, these had attained temper¬ 
atures of 320 C and 240 C, respectively. After 4 min, 
during which these remained practically unchanged, 
firing was resumed. One hundred and eighteen rounds 
fired over a period of 15 min resulted in temperatures 
of 430 C and 360 C, attained 2 min after the close of 
firing. At rates of fire of 5, 10, 15, and 20 rpm, the 
estimated times for the temperature near the breech 
to rise 500 degrees centigrade were found to be re¬ 
spectively 100, 38, 25 and 18 min. 

For a 75-mm gun, M1897E3, temperature-time 
graphs were obtained by means of thermocouples 12, 
36, and 553^ in. from the muzzle when 300 rounds 
were fired at a mean rate of 16 rpm. 181 Maximum tem¬ 
peratures recorded were almost identical, being320 C, 
328 C, and 328 C. It was calculated that at the ther¬ 
mocouple nearest the breech, rates of 10, 15, and 20 
rpm would result in a temperature of 400 C, in 55, 28 
and 19 min, respectively. 

British Experiments 

The British have conducted a temperature investi¬ 
gation 427 with the quick-firing 3.7-in. gun Mark VI. 
The data are graphed in Figure 10. The firing was in 
bursts of 10 rounds fired automatically; the rate of 
fire during a burst being 20 rpm. However, the bursts 
were fired at irregular intervals, as shown by the 
figure. In all 140 rounds were fired in about an hour. 

A spring-loaded thermocouple was used at the muz¬ 
zle; for chamber measurements two spring-mounted 
thermocouples were fastened to a steel tube; this was 
inserted so that contact was made by the thermo¬ 
couples with the wall at the commencement of rifling 
and at the neck of the cartridge case. Temperatures 
were taken at the end of each burst. The relatively 
rapid heating and cooling at the muzzle are clearly 
shown in the figure. The air temperature was 10 C, 
and a high cold wind assisted in cooling the muzzle 
section. It must be noted that the opening of the 
breech for the insertion of the thermocouples had a 
marked cooling effect. Maximum temperatures were 


£200 




j\i 

\ — 

-MUZZLE 

-C OF R 42 IN. FROM 


A / 

|'o 

,\i I 
1 & 

\ 

BREECH FA( 
-CHAMBER 3 
BREECH FAi 
TIME OF FIRI 
OF 10 ROUND 

:e 

2 A IN. FROM 
CE 

NG A SERIES 
iS 

J 

_ 

\/ 

Q 

Z 

UJ 

/ 

/ 

./ 

\ 

\ 

\ 

\ 


l> /: " 

1_1_1 

_it n_ij 

t Iff ft 


-^ 

- 

\ 


TIME IN MINUTES 


Figure 10. Temperature measurements in 3.7-in. AA 
gun, Mk VI. (From Proceedings of the Ordnance Board 
No. 27,604.) 


333 C at the muzzle, 165 C at the commencement of 
rifling, and 73 C in the second chamber position. 

These experimental results have been compared 401 
with computations made by Hicks and Thornhill by 
their method outlined in Section 5.4.1, using an as¬ 
sumed average rate of fire of 2j/£ rpm. Since the early 
firings were at a slower, and the later ones at a faster 
average rate than this, the calculated temperatures 
for the muzzle are high from the second burst till near 
the end of firing, when they agree well with the meas¬ 
urements. Computed and measured temperatures 
near the commencement of rifling agree well through¬ 
out. 


U. S. Naval Guns 

This theory was extended, 401 although with neces¬ 
sary qualifications, to the estimation of temperatures 
in a proposed U. S. Naval 3-in./70-cal. gun to fire 
at a rate of 90 rpm. Upper and lower bounds for heat 
input were assumed. The resulting estimated temper¬ 
ature-time-rounds relations are shown in Figure 11. 
It is to be noted that temperatures near the origin of 
rifling are calculated for operation with and without 
a cooling jacket. The results indicate hypothetical 
mean barrel temperatures, and give no idea of the 
temperature gradients across a barrel section. It was 
tentatively concluded that general softening of the 
bore might occur near the muzzle, but would not be 
likely near the origin of rifling; and that it is unlikely 
that more than two or three bursts of 75 rounds each 
could be fired before the condemning limit of wear 
would be reached. 

A more complete study of the same 3-in./70 cal. 


CONFIDENTIAL 















120 


HEATING OF GUNS DURING FIRING 


ROUNDS FIRED 



Figure 11. Estimated temperatures in proposed 3-in./ 
70-cal. naval gun. (From report to Bureail of Ordnance, 
Navy Department, from E. P. Hicks and C. K. Thorn¬ 
hill. 401 ) 



Figure 12. Average bore surface temperature of 3-in./ 
70-cal. gun firing at rate of 90 rounds per minute, un¬ 
cooled. (This figure has appeared as Figure 72 in NDRC 
Report No. A-434.) 



Figure 13. Outside surface temperature of 3-in./70- 
cal. gun firing at rate of 90 rounds per minute, uncooled. 
(This figure has appeared as Figure 73 in NDRC Report 
No. A-434.) 

gun was later made at the Geophysical Laboratory 
by methods that did take into account position with¬ 
in the barrel wall (see Section 5.4.2). The principal 
results are shown in Figures 12, 13, and 14. At the 


Hi 



Figure 14. Average bore surface temperature of 3-in. / 
70-cal. gun firing at rate of 90 rounds per minute, water- 
cooled. The origin and midway temperatures are prac¬ 
tically the same as in Figure 12 of Chapter 5. (This fig¬ 
ure has appeared as Figure 75 in NDRC Report No. A- 
434.) 


muzzle, where the wall is thin, the average tempera¬ 
ture for an uncooled gun is little larger at the bore 
than at the outer surface (Figures 12, 13). Therefore 
at either position the temperature is nearly equal to 
the mean muzzle temperature, and the agreement 
with Figure 11 is good. But toward the breech, where 
the wall is thicker, the mean temperature across a sec¬ 
tion does not give the essential picture. At the end of 
90 rounds fired in 1 min, the average temperature at 
the bore near the origin of rifling reaches 650 C, while 
at the outer surface it has not begun to rise. It follows 
that external water cooling has no effect at the origin 
during this time, contrary to what appears in Figure 11. 

The Bureau of Ordnance of the Navy Department 
has measured external temperatures near the muzzle 
of 5-in. guns during periods of rapid fire. For a 5-in./38- 
cal. gun, fired 151 rounds in 12% min, a rise of tem¬ 
perature amounting to 375 degrees centigrade was 
observed. For a 3-in./50-cal. gun the temperature 
rose 275 degrees centigrade after 98 rounds had been 
fired in about 16 min. The rise of temperature during 
the latter firing is shown in Figure 6 in comparison 
with calculated values discussed at the end of Section 
5.4.2. Similar calculations have also been made for 
the same gun fired at 40 and 45 rounds per min- 
ute 99 *! 06 and for 3-in. and 90-mm guns fired with dif¬ 
ferent powders at 20 rounds per minute. 


56 RESULTS OF HIGH TEMPERATURES 
561 Erosion 

The temperature of the bore surface is the general 
regulator of the changes there which, taken together, 
constitute erosion, as is discussed in detail in Chapter 
13. One of those changes is liquefaction of the surface. 
In this connection it is important to keep in mind that 


CONFIDENTIAL 





















RESULTS OF HIGH TEMPERATURES 


121 


the liquefied material on the surface of a steel barrel 
is a reaction product having a fusion temperature con¬ 
siderably lower than that of steel, as is brought out in 
Section 12.5.3. Therefore there is no essential conflict 
between the observations summarized in Section 12.6 
and the calculations of bore surface temperatures that 
had led to the conclusion 48 that “the maximum tem¬ 
perature of the bore for all small guns (at conven¬ 
tional velocities) remains definitely below the melting 
point of steel, even with the hot FNH-M2 powder 
and under conditions of rapid firing.” 

5,6 2 Danger of Cook-Offs 

A cook-off mayjbe defined as an explosion of the 
fuze, projectile filler, or propellant that results from a 
high temperature reached by a round after standing 
for some time in the chamber of a hot gun. 

Experimental evidence regarding the danger of 
cook-offs has been gathered from two sources—tests 
definitely set up to determine conditions under which 
cook-offs occur, and studies of the phenomenon as it 
took place fortuitously in the course of experiments 
designed for other purposes or under actual service 
conditions. Temperature measurements in the various 
trials reported were made by thermocouples attached 
to the outer surface of the guns. 

Tests on the caliber .30 Colt-Browning MG firing 
ball, Ml ammunition, and on the caliber .50 Brown¬ 
ing MG were conducted in England. 381 For air-cooled 
guns, cook-offs of detonator, filling, or propellant 
were shown to result after long bursts. Bursts of 100 
to 200 rounds caused cook-offs of the fuze and shell¬ 
filling; bursts of 250 and more rounds led to cook-offs 
of all three components. In water-cooled guns no 
cook-offs resulted. 

In experiments 526 with cooling of the caliber .50 
machine gun, gun temperature against time to cook¬ 
off were plotted for 14 observed cases. This showed 
that the time decreases with temperature. No cook¬ 
offs were observed at temperatures below 900 F, 
which was considered a critical temperature. 

In a study of the 20-mm Hispano Mark Vgun, tests 
were again made for cook-offs of the various compo¬ 
nents. Several different fillings and propellants were 
tested. Some cases of explosions in the lips of the belt 
feed mechanism were reported. As a result of these 
trials, although bursts of 175 rounds caused no cook¬ 
offs, a maximum of 150 rounds was recommended for 
complete safety. However, in a test of 20-mm A A 
guns conducted by the Navy, Bureau of Ordnance, a 


cook-off occurred after the firing of only 120 rounds. 
This may have been due to a very high rate of fire. 

Results also vary with 40-mm guns. The lowest 
reported number of rounds causing a cook-off of U. S. 
ammunition, starting with a cold gun, is 94. Firings in 
England using British guns and ammunition, have 
been reported to give cook-offs 3 min after the firing 
of 50 rounds, with an initially cold gun. As to water- 
cooled guns, a Bureau of Ordnance letter states “it is 
believed that relatively small danger of a cook-off 
exists in the water-cooled gun when the cooling system 
is operating properly.” 

The British have also conducted trials for propel¬ 
lant cook-offs in the 3.7-in. gun, Mark VI, and the 
quick-firing 3.7-in. A A guns, Marks I and III. It was 
concluded that one of the cooler propellants would 
not cook-off in a round left in the gun after a series of 
firings at a rate of 20 rounds per minute and that the 
hotter ones would be unlikely to do so under most 
circumstances. 

Relatively little information is available on cook¬ 
offs in larger guns, most of this having been gleaned 
from field experience or in tests set up for other pur¬ 
poses. Taking into consideration the dependence of 
heat input on caliber, the following rule-of-thumb for 
heating has been stated by British authorities: 544 

The number of rounds of continuous fire with full charges 
that will bring guns to the conditions described as “hot” will 
vary with rate of fire and may be taken to be approximately 
as follows: 

6-inch and above 80 rounds 

Below 6-inch 30 rounds 

If a gun has not reached this critical condition no special pre¬ 
caution need be observed in leaving a round chambered. 

5.6.3 Eff ec t 0 f Temperature on Ballistics 

Accuracy Drop from Expansion of Bore 

When a gun barrel wall reaches high temperatures 
as the result of rapid fire, the metal expands. If this 
expansion is sufficiently great, the engagement of the 
rotating band of the projectile with the rifling of the 
bore diminishes until proper spin is no longer im¬ 
parted to the projectile, when tumbling begins. 

It is interesting to compare the ballistic behavior of 
a barrel with the expansion computed from a meas¬ 
ured rise of temperature. This has been done for a 
caliber .50 aircraft barrel with a 9-in. stellite liner. 106 
During the firing of a 425-round burst, tumbling be¬ 
came persistent after about round 350. At this time the 
outer surface temperature at a distance of 24.0 in. 


CONFIDENTIAL 



122 


HEATING OF GUNS DURING FIRING 


from the breech was 735 C, and the temperature 0.14 
in. from the bore surface was 765 C. Assuming an 
average barrel temperature of 750 C, and a coefficient 
of linear thermal expansion of 16X10 -6 , the increase 
in bore diameter amounted to 0.006 in. Since the 
depth of rifling in this barrel is only 0.005 in., the 
amount of spin imparted to the bullet would be small 
because of “skidding,” an effect that is illustrated in 
Figure 12 of Chapter 27. 

At less extreme temperatures there is still some 
loss of spin, accompanied by balloting and increasing 
yaw within the gun, resulting in greater external yaw 
and loss of accuracy. This effect in a caliber .50 ma¬ 
chine gun barrel has been counteracted to a large ex¬ 
tent by using a “choked muzzle” obtained by means 
of chromium plate of tapering thickness, as described 
in Chapter 23. 

Pressure and Velocity Changes from Heated 
Propellant 

In ballistic tables, pressure and muzzle velocity are 
calculated for a powder temperature of 70 F (21 C). It 
has long been known, however, that an increase in 
temperature increases the potential and the burning 



Figure 15. Average velocity and pressure vs temperature 
nical Division; High and low temperature ballistic research , Is 


rate of the powder (Section 3.3.3), giving a greater 
muzzle velocity. 509 515 Pressures are also greater. Since 
it w r as recognized that in service a powder might be 
fired at ambient temperatures varying from — 50 F to 
+ 150 F, this problem was studied by the Army 
Ordnance Department. 

Tests were made 286 with FNH-M1 powder manu¬ 
factured for 3-in. guns fired in a 76-mm gun, Ml. The 
special interest in this experiment was in the low r 
temperatures, since service experience had shown 
that dangerous sporadic pressures may arise under 
conditions of extreme cold. However, the tests at 0 F 
and above, w r hile considered “satisfactory and nor¬ 
mal,” showed a definite linear trend of increased 
pressure with rising temperature, shown in Figure 
15. In the course of these tests a procedure w r as devel¬ 
oped and used for ten 76-mm guns, whereby the 115 
per cent pressure proof-firing w^as performed, using 
rounds heated to 135 F rather than by increasing the 
charge. If we assume that the linear trend exhibited 
in Figure 15 would continue, a powder temperature of 
265 F (129 C) would yield a dangerously high pressure 
of 145 per cent normal. This temperature might well 
be reached by a round left for a considerable time in 
the chamber of a hot gun. 



30,000 35,000 40,000 45,000 50,000 55,000 

PRESSURE IN PSI (COPPER) 


for 76-mm gun, Ml. (By courtesy of War Department Tech- 
t Progress Report. 286 ) 


CONFIDENTIAL 






























RESULTS OF HIGH TEMPERATURES 


123 


These experiments were continued with a 37-mm 
gun, T28 (ballistically the same as the M3) and with a 
20-mm gun, AN-M2. j In the latter case, using IMR 
powder, it was demonstrated that the change in pres¬ 
sure was approximately linear and equal to 49 psi per 
degree Fahrenheit change in temperature. For the 
37-mm gun tests were carried out with both FNH-M5 
and FNH-M2 powders. Equations (37) and (38) were 
developed in the two cases for the change in pressure 
in pounds per square inch as a function of Fahrenheit 
temperature P. 

( Pt ~ Pes) = - 3421 + 42.5P 

-.008P 2 + .0018P 3 (37) 
where P 68 = 45,300 psi when P = 68 F for FNH-M5 

(Pt - P 68 ) = - 2584 + 22.8P 

+ .066P 2 + .0023 P 3 (38) 

where P 68 = 43,100 psi when P = 68 F for FNH-M2. 

For muzzle velocity in the 76-mm gun, the indi¬ 
cated increase per degree Fahrenheit was shown to be 
approximately 1 fps. For the 20-mm gun, the velocit}^ 
change was a linear function of the temperature equal 
to 1.2 fps per degree Fahrenheit. In the 37-mm gun, 
the trends were found to be nonlinear, as in the case 
of the pressure, equations (39) and (40) having been 
developed to express the data. 

(V T - V«s) = ~ 81.7 + 1.416P 

- .00599P 2 + .0000417 P 3 (39) 
with V 6 8 = 2,894 fps for P = 68 F for FNH-M5. 

(V T - V 68 ) = - 59.1 + .89P 

- .00304P 2 + .000040P 3 (40) 
with F 68 = 2,842 for P = 68 F for FNH-M2. 

5 6 4 Weakening of the Barrel Wall 

That high temperature attained in a gun can result 
in weakening of the barrel wall has long been well 
known. The sequence of events in long bursts of auto¬ 
matic fire in a steel barrel appears to be as follows. 
During the early rounds the conduction from the bore 
surface is rapid; the barrel as a whole heats slowly 
with slight and generally favorable effects on ballis¬ 
tics. After the mass temperature of the barrel be¬ 
comes relatively high, thermal expansion permits 
some balloting of the projectile. At the same time the 
heated bore surface begins to be austenitized and the 


j Private communication from B. E. Anderson, Office of the 
Chief of Ordnance. 


lands to be swaged. The swaging combined with 
further thermal expansion weakens the engagement 
of the rifling, increasing internal yaw of the projectile 
and making obturation less effective. Origin erosion 
with consequent blow-by of gases results in reduced 
muzzle velocity and still higher bore surface tem¬ 
perature. 

The process continues and the bullets pursue erratic 
courses in the hot bore. If a short section at the muz 
zle has a small enough diameter and sharp rifling (as 
in a barrel with choked muzzle chromium plate), the 
bullet may receive sufficient spin to be stable after 



Figure 16. Caliber .50 aircraft barrel distorted during 
firing. Note that bulging is confined essentially to one 
plane; the two photographs were taken at right angles 
to each other. 


CONFIDENTIAL 





124 


HEATING OF GUNS DURING FIRING 


leaving the barrel. Otherwise tumbling occurs. Fi¬ 
nally, at very high temperatures the tensile strength 
of the gun steel fails, the barrel warps and the gun 
jams; in extreme cases the bullet plows into, or even 
through, the barrel wall, as is illustrated in Figure 16. 
In the case of a stellite-lined barrel, origin erosion is 
inappreciable and tumbling (keyholing) begins rather 
suddenly, with no great drop in muzzle velocity. 

In the case of larger guns, this problem of the effect 
of temperature on yield strength has received careful 
consideration. In rapid-fire tests of a 3-in. AA gun, 
calculations were made to determine maximum tem¬ 
peratures permissible at various distances from the 
breech. 179 These were determined from the expected 
pressures and the yield points of the gun-steel at 
elevated temperatures. On the basis of these compu¬ 
tations, it was decided to have the crew take cover 
when a welded thermocouple 5.21 in. from the muzzle 
indicated an external barrel temperature of 250 C. 
The actual temperature there at which bursting 
might have been expected was approximately 475 C. 
Although the former temperature was maintained for 
a considerable length of time, the latter was not 
reached during the firing, the highest temperature 
recorded by this thermocouple having been 360 C. 
This temperature was attained 2 min after the firing 
of 118 rounds in 15 min; previous firing had already 
preheated the gun to 245 C. 

Similar calculations were made for the 105-mm AA 
gun, Ml. Dangerous external temperatures ranged 
from 405 C at 165 in. to 565 C at 40 in. from the muz¬ 
zle. This gun has a liner with a clearance tapering 
from 0.007 in. at the muzzle to 0.003 in. at the breech. 
Most of this space is filled with mixed grease and 
powdered graphite when the liner is inserted. 

A number of mechanical effects of raised tempera¬ 
tures in guns have been listed by the British Gun De¬ 
sign Committee. 333 A first statement is “The temper¬ 
atures attained (a maximum external value of 322 C) 
would not permanently affect the yield point of the 
ordinary gun steels because they are far below the 
tempering temperatures employed, but the yield 
point of the steels when stressed at the raised tem¬ 
perature would be lower than at normal temper¬ 
atures.’ ; 

It was also remarked that uniform heating would 
not substantially reduce internal stresses due to auto- 
frettage or to shrinkage, but “The heating is, of 
course, not uniform and a steep temperature gradient 
might have more pronounced effects on stress distri¬ 
bution. Temperature gradients through the wall may 


be very irregular in guns which have been cooled 
after rapid fire.” 

In conclusion there was considered the possibility 
of longitudinal stress. It was pointed out that a tem¬ 
perature difference of 100 degrees centigrade between 
the inner and outer tubes of a built up gun might lead 
to protrusion of the liner at the muzzle. This, togeth¬ 
er with circumferential cracking, has been observed 
in 8-in. howitzers, Mark VII; they were, however, 
considered to be poorly designed. It has also been 
considered in connection with the design of a 16-in. 
gun. 429 It is stated that a difference of 400 degrees 
Fahrenheit may cause a loose liner to protrude as 
much as 2 in. This may cause the liner to bell-mouth 
thus affecting accuracy. It may later prove difficult 
to remove the liner. Additionally, appreciable hoop 
stresses may be caused by the liner being wedged into 
the taper seating by reason of its longitudinal exten¬ 
sion and diametrical expansion. 

5 7 COOLING OF GUNS k 

571 Rates of Cooling 

Having examined the rates and sources of heat in¬ 
put to the gun bore, and considered the distribution 
of the heat in the barrel wall, we turn our attention to 
the corresponding cooling conditions; in particular, 
we shall look for methods by which the temperatures 
may be held within limits which will prevent the 
occurrence of the most serious results of overheating. 

There is no doubt that some heat loss takes place 
from the bore surface; the opening of the breech of 
the gun causes some forced convection there, and 
ejection of the projectile has some effect at the muzzle. 
But in the case of very rapid fire there is little chance 
for the gases within the gun to cool, and although as 
has been pointed out, 348 there is some reversal of heat 


k Division 1, NDRC did not undertake any extensive ex¬ 
perimental investigation of the cooling of guns, but limited its 
activities largely to interpreting data obtained elsewhere, in an 
effort to apply them to the problem of increasing muzzle 
velocity without decreasing barrel life. In an effort to facilitate 
the crossflow of information it organized, at the instance of the 
Army and Navy, a special Advisory Committee on the Cooling 
of Guns, made up of representatives of the Army Ordnance 
Department, the Air Ordnance Office, the Navy Bureau of 
Ordnance, the Navy Bureau of Aeronautics, two Division 1 
contractors (Geophysical Laboratory and Crane Company), 
and Division 1 staff. This committee held three meetings dur¬ 
ing 1945 and two during 1946. Much of the material in this 
section was presented and discussed at those meetings. 


CONFIDENTIAL 






COOLING OF GUNS 


125 



A STILL QUIET AIR-OUTSIDE RANGE 



* 


1 THREE l" DIA JET TUBES-3/8 "DIa\jETS 

CLOSE TO BARREL JACKET 


Q 

J 



T7 —U u O-O—O—O-C7—o-O-1_ 

ooooooooooP 


THREE I DIA JET TUBES 
l" DISTANT FROM JACKET 




RADIALLY OPPOSITE 




5 a g g .y u 

TWO JET TUBES-OPPOSITE 
DIRECTED TANGENTIALLY 






NO SLEEVE 



LONG GUIDE TYPE 


0 

M 


SINGLE I DIA JET TUBE 


a (°C 


D 


\o-0^o”o^oYb\o“(^o\o b 

^ A A A - A, , A A A /T k ... . .J 


45° ANGULAR JET TUBE 
3/8" JETS 






\\\UJA 



MODIFIED P-51 H GUIDE 
P TYPE COOLER 


Figure 17. Various types of cooling devices tested on ground (by permission of North American Aviation, Inc.; Report 
No. NA-8568, Performance and design criteria for caliber .50 gun barrel coolers for aircraft). 


CONFIDENTIAL 









































































































126 


HEATING OF GUNS DURING FIRING 


flow, this must have very slight effect and we may 
continue to consider the flow as one-directional. 

On the external surface of the gun there is longi¬ 
tudinal variation in temperature; commonly the 
thinner muzzle section heats more rapidly at first, but 
as flow continues through the wall the greater heat 
input near the origin of rifling results in higher tem¬ 
peratures in that region. For our purposes it is con¬ 
venient to consider the whole outer surface as being 
at approximately the same temperature. Cooling then 
takes place by radiation, and by free or forced con¬ 
vection. 

Radiation plays an important part in cooling. It 
has been estimated 187 that “on a windless day, with 
the gun at a temperature of 330 C almost 80% of the 
heat loss is by radiation.” The intensity of radiation 
is given by the Stefan-Boltzmann law, equation (41), 

/ = <r(7V - 7VK, (41) 

in which Tq is the absolute temperature of the gun 
surface, T A is ambient temperature, and e s is the 
emissivity of the barrel surface. If I is in calories per 
second per square centimeter, the value of <r is 
1.36 X 10~ 12 . For the example just given, we find 
/ = 0.18 cal/cm 2 . If this rate were maintained, a gun 
with a heat input of 18 cal/cm 2 in one round would 
have its heat dissipated in about 50 sec if we assume 
a wall ratio of 2 and ignore the time of conduction 
through the barrel. Of course the radiation rate drops 
rapidly with temperature; but this emphasizes the im¬ 
portance of cooling by radiation at high temperatures. 

Cooling by convection is fundamentally exponen¬ 
tial. Since the cooling rate is proportional to the dif¬ 
ference between surface and ambient temperature, as 
given b}' equation (42\ 

d J = - k( T rj - T a ), (42) 

we obtain the temperature-time equation (43): 


Since cooling actually results from the two causes 
acting simultaneously, various empirical relations 
have been adopted to represent the situation. One 
assumption 187 was that cooling is proportional to the 
1.23 power of the temperature difference. Another 
investigator 181 assumed the power to be 5/4 and 
made experimental determination of the proportion¬ 
ality factor for the 75-mm gun, pointing out that this 
is dependent on wind velocity. The resulting graphs 


indicated cooling of approximately 280 centigrade 
degrees in 80 min. The temperature at the thermo¬ 
couple nearest the breech did not drop this rapidly, 
indicating greater heat storage in that section. In¬ 
crease in thickness of the barrel wall giving greater 
heat storage has long been known to diminish max¬ 
imum temperature. The limitation of the method is 
obvious; the added weight is particularly undesirable 
in aircraft guns. 

5 7 2 Methods of Cooling 

Air blasts. The noticeably improved cooling rate 
occurring during firings made in strong winds sug¬ 
gested the desirability of using artificial air blasts. 
The weight of other types of cooling equipment, com¬ 
bined with the obvious possibilities of forced convec¬ 
tion in flight, made this the natural method for air¬ 
craft. Extensive studies of this type of cooling were 
made both in Britain and in this country. 

The cooling system 302 ’ 411 ’ 419 ’ 420 - 430 developed in Brit¬ 
ain by the Naval Air Fighting Development Unit 
consists of a 23 ^-in. diameter steel tube surrounding 
the barrel of the caliber .50 gun and open at both 
ends. An auxiliary tube 1 in. in diameter enters the 
rear portion of the tube and extends to the outside of 
the lower wing surface. One assembly weighs approx¬ 
imately 5 lb. The assemblies have been adapted for 
use with a number of types of fighter planes. When a 
firing cycle of 3-sec bursts at 1-min intervals is used, 
the N.A.F.D.U. cooling tube doubles the accuracy 
and velocity life of chromium-plated barrels and gives 
marked improvement in standard unplated barrels. 
It will not, however, be of benefit if continuous bursts 
of 300 rounds are fired. 

A comprehensive program of tests for performance 
of a variety of cooling devices was carried out 525 - 526 for 
the Army Air Forces at North American Aviation, 
Inc. Thermocouples were fastened to the barrels at 
five points spaced at 5-in. intervals along the barrel, 
beginning at a point 5 in. from the face of the receiver. 
Temperatures at these positions were graphed for 
various times during the firings. These records were 
made for a number of firing schedules with the several 
cooling installations. The cooling devices tested are 
shown diagrammatically in Figure 17. 

The second of the thermocouple positions was 
adopted as a temperature index for performance; 
occasionally more than one thermocouple were placed 
there. The performance factor for a given cooler rep¬ 
resents the number of rounds that can be fired each 


CONFIDENTIAL 




COOLING OF GUNS 


127 


minute through a barrel that has already reached a 
temperature of about 900 F, without further raising 
its temperature appreciably. Numerical and graph¬ 
ical methods were established for determining this 
point. Some typical values of the performance factor 
are given in Table 8. An empirical linear relation be¬ 
tween this factor F and the length I of the permis¬ 
sible initial burst is given by equation (44). 

I = 125 + 0.65 F (44) 

A single jet cooler was decided upon as most effective. 

Water Jackets. Cooling of gun barrels b}^ means of 
water circulating through a jacket over the barrel 
was considered and adopted many years ago. Where 
this type of installation is feasible, it has proved to be 
very effective. Water-cooled machine guns have oper¬ 
ated under severe schedules without overheating. In 
some cases the heat of vaporization is utilized by per¬ 
mitting boiling off of the water. 

The objections to water-cooling are weight, size of 

Table 8. Performance factors of air-cooling devices for 
caliber .50 aircraft machine gun barrels fired 200-round 
bursts in still air on the ground.* 


Per- 

Type of cooler formance 

(See Figure 17) factor f 


A. (No cooling) 10 

B. Full muff cooler, at center 19 

C. Full muff cooler, inboard (Cooler against gun 

trunnion collar) 14 

D. Half muff cooler, inboard (Air inlet near body) 17 

E. Short muff cooler and sleeve 64 

F. Short muff cooler without sleeve 37 

G. Long guide type cooler 40 

H. Short guide type cooler 51 

I. Triple jet tube cooler, close to barrel jacket 48 

J. Triple jet tube cooler, 1-in. distant from jacket 32 

K. Two jet tubes, radially opposite 41 

L. Two jet tubes opposite, directed tangentially 42 

M. Single jet tube 

Jets aimed at barrel jacket holes 55 

Jets aimed at solid jacket 28 

Jets aimed between holes in jacket 53 

Jets aimed at holes, with P-51H guide 55 

N. 45° angular jet tube 

Jets aimed opposite thermocouple 45 

Jets aimed at thermocouple 31 

O. Manifold type cooling tube 

Touching jacket, 51 in. from collar 56 

Touching jacket, 3j in. from collar 49 

One inch from jacket, 11 in. from collar 20 

One-half inch from jacket, 31 in. from collar 33 

P. Modified P-51H guide type cooler 29 


* This table has been taken by permission from the summary on pages 
59 and 60 of Report NA-8568, North American Aviation, Inc. 525 

t The “Performance Factor” is the number of rounds that can be fired 
each minute through a barrel that has already reached a temperature of 
about 900 F, without raising its temperature appreciably. 


installation necessary, and under some conditions, 
difficulty in securing water for the purpose. 

Tests have recently been made at the Naval Prov¬ 
ing Ground on the cooling of 40-mm guns by this 
method, as referred to in Section 5.5.2. In addition, as 
mentioned in Section 5.6.2, service reports indicate 
no dangerously high temperatures in these guns. 

In the 5-in./38-cal. gun tests mentioned in Section 
5.5.3, the effect of external water-cooling by hoses 
was tried. After the close of firing, a projectile was in¬ 
serted in the breech. Temperatures under the band 
and under the bourrelet rose for about 6 min after the 
water was applied, then fell off as it continued to play. 

Special Coolants. The use of special coolants has 
often been considered. Preliminary tests of a cooling 
device which injects liquefied C0 2 into the chamber of 
a caliber .50 barrel during firing were made at the Geo¬ 
physical Laboratory. 110 One hundred-round bursts, 
with a 2-min interval between bursts, were fired in a 
new steel aircraft barrel, using AP-M2 ammunition. 
The rate of fire averaged 810 rpm. Results showed 
apparent improvement over a similar barrel without 
cooling, but a direct comparison with earlier firings 
was impossible because of three short stoppages in 
the second burst. The amount of C0 2 used during the 
firing of 500 rounds was 4.6 lb. 

Injection Sprays. Cooling of the bore by wet swabs 
is traditional; and it has been suggested that for 
smaller caliber cannon, the tube might be cooled by 
pumping water through the bore between bursts. The 
general problem of utilization of the latent heat of 
vaporization of water has been frequently studied, 
and injection-cooling has now been developed. 

The initial work in this development was done with 
caliber .50 barrels by Purdue University for the Army 
Ordnance Department, using a nonaqueous cooling 
liquid. Although considerable improvement in the 
performance of steel aircraft barrels was found, 294 it 
was not as great as that gained by the use of either a 
stellite liner (Chapter 22) or choked-muzzle chromium 
plate (Chapter 23). Some tests by the Ordnance De¬ 
partment with stellite-lined barrels and chromium- 
plated barrels delivered by Division 1 showed still 
further improvement in the former by the use of this 
device, but the result was not obtained in general 
with chromium-plated barrels. Furthermore, the de¬ 
vice was erratic in its performance 1 and decreased the 
cyclic rate of the gun. 

1 A few tests on special barrels using this cooling device were 
made by Division 1 at the Geophysical Laboratory, 81 ' 110 but 
mechanical difficulties prevented any comprehensive testing. 


CONFIDENTIAL 








128 


HEATING OF GUNS DURING FIRING 


A similar injection-cooling device that was more 
satisfactory mechanically was developed by Purdue 
University 545 for the 90-mm gun, M1A1, and the 155- 
mm gun, Ml. For bores this large it was possible to 
impart rotation to the spray in an attempt to disperse 
any tendency toward “film” conditions. These cool¬ 
ing devices were tested at Aberdeen Proving Ground. 

The nozzle was manually positioned. Water was 
injected at a gauge pressure of 200 psi. For the 90-mm 


gun a weight of 0.47 lb of water proved satisfactory 
for a rate of fire of 10 rounds per minute. This held 
the temperature to 260 F with very slight tempera¬ 
ture gradient. For a higher cyclic rate, 0.60 lb of wa¬ 
ter seemed preferable. For the 155-mm gun, 1 lb of 
water was injected in a 2-sec cycle. At the time of 
writing the tests were being continued. They indi¬ 
cated favorable results for this method of cooling as a 
means of reducing erosion. 




CONFIDENTIAL 




Chapter 6 

BORE FRICTION 

By W illiam S. Benedict 


INTRODUCTION 

Nature and Importance of 
Bore Friction 


T he ideal gun would be one in which all the avail¬ 
able energy of the powder was transformed into 
kinetic energy of the projectile. Unhappily, as pointed 
out in the discussion of the general problem of hyper¬ 
velocity (Chapter 1) and more particularly in Section 
3.5, actual guns fall far short of the ideal, and even 
short of the more attainable goal of an efficient ther¬ 
modynamic engine, in several respects. One principal 
source of inefficiency, thermal losses to the bore, has 
been discussed in the preceding chapter. Another 
most important cause of diminished performance in 
new guns, a factor of some importance in their ero¬ 
sion, and a prime factor in their erratic behavior, is 
the subject of the present chapter, bore friction. 

Under the heading of bore friction it is customary 
to include all the forces that oppose the acceleration 
of the projectile, counter to the force exerted on the 
base of the projectile by the pressure of the powder 
gas. We will express the friction as the force per unit 
area that resists the accelerating pressure on the base 
of the projectile, and denote it by P r , usually in units 
of thousands of pounds per square inch, klb/in. 2 . If 
the inertial mass of the projectile is m, the cross-sec¬ 
tional area of the bore A, X , the coordinate of the 
projectile relative to the gun, and Px the pressure on 
the base of the projectile when it is at X, equation 
(1) is the fundamental one by which we define the 
friction. 


P r = 


d 2 X 


A dt 2 


+ P 


x, 


( 1 ) 


The largest portion of this chapter is concerned 
with the methods of determining the accelerating 
pressure, the base pressure, and their difference, 
which is the resisting pressure or bore friction; and in 
describing the results obtained in measuring these 
quantities in the ballistic firings carried out as part 
of the program of Division 1 at Carderock (Chapter 


a Geophysical Laboratory, Carnegie Institution of Washing¬ 
ton, and National Bureau of Standards. 


4). As a preliminary, we shall discuss in general terms 
the nature of the various forces that may be expected 
to contribute to the resisting pressure. 


6,1,2 Component Forces Appearing as 
Bore Friction 342 

Engraving Forces 

At the commencement of motion, in guns emploj^- 
ing conventional banded projectiles, the resisting 
pressure will be high compared to the accelerating 
pressure. At the very start, with fixed ammunition, 
the projectile must be released from the case into 
which it is crimped. Then follows the engraving peri¬ 
od. The force required to engrave the rotating band, 
as determined from “static” tests in which the pro¬ 
jectile is slowly forced through the rifling in a testing 
machine, ranges from 2 to 20 klb/in. 2 . The force is re¬ 
quired to “engrave” the soft metal of the band by the 
process described in Section 7.3.5 and to overcome 
surface-to-surface friction between the band and the 
bore surface. In guns in which the projectile is ram¬ 
med home against the forcing cone these engraving 
resisting forces will be operative at the very start, so 
that the velocity of the projectile is zero as it enters 
the forcing cone, and the starting friction during fir¬ 
ing should be equivalent to that determined stati¬ 
cally; in most guns with fixed ammunition, however, 
the projectile will have traveled on the order of a 
tenth of a caliber and have acquired an appreciable 
velocity before engraving begins. In all cases the ve¬ 
locity will increase during engraving. With increased 
velocity the resistance to plastic deformation will de¬ 
crease, and hence the engraving friction will fall be¬ 
low that found statically. 


Band-to-Bore Friction 

After the deformation of the band to fit the grooves 
of the rifling is complete, the passive resistance gen¬ 
erated at the band decreases greatly, being now due 
principally to surface friction. (In guns with increas¬ 
ing twist engraving continues throughout the travel; 
in most guns it is virtually complete after the rear 


CONFIDENTIAL 


129 



130 


BORE FRICTION 


of the band has entered the region of constant land 
diameter.) The surface component of friction P r8 is 
proportional to Pb, the pressure of contact of the 
surfaces (band pressure), as expressed by 


P 


rs 


4 fbPb 
D ’ 


( 2 ) 


where b is the length of band, D the tube diameter, 
and / the coefficient of friction. As discussed in Chap¬ 
ter 7, Pb depends only to a slight extent on the pow¬ 
der pressure; it is more a function of the relative 
dimensions of band and tube, and is nearly constant 
over the length of the gun, decreasing somewhat to¬ 
wards the muzzle. For most bands Pb is of the order 
of 50 to 70 klb/in. 2 . The coefficient of friction/, as is 
well known, 517 tends to decrease rapidly below its 
static value (which for copper-steel is of the order of 
0.1) with increasing velocity; it should also depend 
on the condition of the bore and band, being de¬ 
creased by lubrication of the bore, and decreasing 
further if a surface layer of the band should melt. The 
theory of the heating of rotating bands 45 indicates 
that melting of the band surface occurs in many guns; 
this is confirmed by the known facts of coppering of 
the bore surface (Section 10.5.4) and by direct meas¬ 
urements 72 of the interface temperature. Thus, / and 
P r8 will be comparatively low for much of the travel; 
the latter probably below 1 klb/in. 2 . 


Body Engraving 

As the projectile acquires velocity, however, other 
factors will enter to increase P r . With increasing wear 
on the band, and especially in partially eroded guns, 
the projectile cannot remain centered in the tube. 
The bourrelet will come in contact with the bore at 
irregular intervals and with varying intensity. In ex¬ 
treme cases this leads to “body engraving/’ wherein 
impressions of the rifling are found to considerable 
depth. (The relation of this phenomenon to muzzle 
erosion and gun performance is discussed in Section 
10.4.10.) Such bourrelet-bore and body-bore contacts 
result in contributions to P r , whose magnitude cannot 
be estimated in advance, but which may vary consid¬ 
erably in time and intensity from round to round. 

Back Pressure of Gas 

As the projectile moves down the tube, it com¬ 
presses the gas ahead of it, both the originally present 
atmosphere and any powder gases that may have 


leaked past. The compressed gas will flow out of the 
muzzle, but its rate of flow cannot exceed the velocity 
of sound (1,100 fps when unheated by compression). 
Hence after the projectile velocity exceeds 1,100 fps 
the back pressure due to the gas becomes increasingly 
important; for an average gun, with a muzzle velo¬ 
city of 2,500-3,000 fps, the resisting pressure due to 
compressed gas at ejection may be of the order of sev¬ 
eral hundred pounds per square inch. 

613 Typical Friction Curves 

In most systems of interior ballistics (Section 3.2.2) 
it is found convenient to divide P r into two portions. 
The relatively large resisting pressure during engrav¬ 
ing is set equal to the starting pressure Po ; it is as¬ 
sumed that no motion of the projectile occurs until 
Px = Po- Thereafter it is assumed that P r is a con¬ 
stant fraction c of Px - It is apparent from the discus¬ 
sion given that this represents a considerable over¬ 
simplification of the true course of the friction-time 
or friction-travel curve. The situation is illustrated 
graphically in Figure 1. A typical pressure-time and 
pressure-travel curve for the 3-in. gun are there pre¬ 
sented. The friction, as determined by methods which 
are described later, is seen to follow the general course 
outlined in the preceding discussion. When plotted as 
a function of travel, the high engraving friction is fol¬ 
lowed by a rapid decrease to a much lower level. The 
conventional friction, with P 0 = 4 klb/in. 2 and c — 
0.04, is plotted as the dashed curves. It is seen that 
while there is considerable difference in the friction¬ 
time curves, the areas under the friction-travel curves, 
which determines the energy expended in overcoming 
friction, as given by 

Ef — A j P r dX, (3) 

are not greatly in divergence. It is further to be noted 
that except during engraving the friction is very 
small relative to the pressure, and that therefore, in 
order to determine it with any degree of accuracy, all 
the experimental factors entering into its calculation 
must be known with great precision. 


6 2 METHODS OF DETERMINING FRICTION 

621 General Discussion of 

Available Methods 

Having sketched the general features of the prob¬ 
lem of bore friction, we now discuss the methods by 


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METHODS OF DETERMINING FRICTION 


131 



10 7.5 5 2.5 0 

X-TRAVEL TO EJECTION IN FEET 


Figure 1 . Typical curves of pressure and friction versus time and travel, 3-in. gun. 


which it may be determined. The most satisfactory 
method would be by direct application of equation 
(1), measuring the base pressure, the acceleration, or 
their difference, by suitably loaded gauges mounted 
in the projectile. Such measurements have been at¬ 
tempted 46 but reliable results have not been attained. 
Not quite so direct, but apparently capable of greater 
precision, are other methods based on equation (1), in 


which the base pressure is inferred from pressure 
measurements made by gauges located at several 
positions along the bore of the gun, and the accelera¬ 
tion is determined from differentiations of a curve by 
locating the projectile with great accuracy as a func¬ 
tion of the time. Finally there are indirect methods, 
by which the friction, possibly averaged over a por¬ 
tion of the gun, is inferred from other measurements, 


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132 


BORF FRICTION 


such as the heat input to the bore surface or the 
strains in the tube wall. In using these indirect meth¬ 
ods a distinction must be drawn between nonballistic 
conditions, in which the frictional forces may be iso¬ 
lated for study, and ballistic conditions; only in the 
latter are the results truly meaningful for the firing 
process, but it is extremely difficult to separate the 
observable results due to friction from those due to 
the action of the powder. 

6 2 2 Indirect Methods 

Static Push Tests and Strain Measurements 

The study of band pressure and related stresses is 
described in more detail in Chapter 7, and need only 
be mentioned at this point. The quasi-static push 
tests , 64 ’ 114,115 of 37-mm and 75-mm projectiles of vari¬ 
ous band designs are of fundamental importance in 
relating the magnitude of the engraving friction to 
the band dimensions and design, but give only an 
upper limit for the actual values of P r to be encoun¬ 
tered in firings. The mathematical theory 13,17,58 of the 
stresses in a gun tube during firing contains relations 
from which the friction may be deduced from simul¬ 
taneous measurements of the axial and tangential 
strain in the barrel as is described in Section 7.1.4. 
The friction is, however, a second-order effect and no 
estimates of its magnitude of any accuracy have as 
yet been obtained from available strain data. 

Falling Weight Tests 

Measurements 72 of the temperature rise in the bar¬ 
rel and at the band-bore interface have been made 
when standard ball M2 bullets were forced through 
sections of caliber .50 barrels. Both nonballistic tests, 
in which a falling weight and piston arrangement 
forced the bullet through the bore at velocities up to 
60 fps, and ballistic firings, with standard and re¬ 
duced charges of IMR powder, were performed as de¬ 
scribed in Section 5.2.2. Their principal aim was to 
evaluate the contribution of the frictional heat to the 
observed heat input in caliber .50 guns; but it also 
appears possible to determine the resisting pressure 
from experiments of this type, as may be seen from 
the following considerations, based on a report 45 on 
the theory of the heating of rotating bands. 

The energy of friction E f , defined by equation (3), 
will appear as heat generated at the interface, to the 
extent that P r represents bore-band friction. It will 


flow into the bore (which is cold at the moment the 
band reaches it) at a rate dQfb/dt and into the band 
(which rapidly heats up) at a rate dQ fp /dt. The tem¬ 
perature of the interface or band surface, T p , de¬ 
pends on these rates and the thermal constants and 
history of both surfaces. Considering the interface to 
be a smooth cylinder, of diameter D and band length 
b, the rate of heat production per unit area of inter¬ 
face, 

dQ f = dEf = APrV = DPrV 
dt 7 rDb dt 7 rDb 4 b 

_ dQfb , dQf P /j\ 

“ dt “*■ dt ' V ; 

The amount of heat produced at any point equals the 
rate times the contact time, b/V. In the customary 
English units (Q in cal/in. 2 -sec; P r in klb/in. 2 ), this is 

Q f = Q f b + Q fP = lMGDPr. (5) 

If the friction is not produced uniformly at the inter¬ 
face (it will be greater along the top and driving sur¬ 
face of the lands than in the grooves in all guns), 
equation (4) must be modified, but equation (5) will 
hold. Assuming the band to be a source of uniform 
temperature T p , we have 



and T p in turn is determined by the integral 

<7 » 

The simultaneous solution of equations (4), ( 6 ), and 
( 7 ), given P r and V as functions of t, permits the deter¬ 
mination of T p , Qfb and Q/ p ; conversely, if T p and V 
are known, P r and the Q’s may be determined. 

In the falling-weight tests, V, T p , and Qfb were de¬ 
termined, the two last being in agreement under the 
theory just sketched. We find that under the condi¬ 
tions of these experiments Qfb and Qf P are nearly 
equal, so that the average 

Pr = 5 x 2 fp 40 - = 3.8 Qf„. At V = 40 fps, Pr = 15 
klb/in. 2 ; and at V = 60 fps, P r = 12 klb/in. 2 . 

In the ballistic firings T p was measured, but since 
the time-travel relations of the projectile are only 
approximate, and pressure data are lacking, it is not 
possible to say more about the friction than that it 
was sufficient to raise the maximum temperature in 


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METHODS OF DETERMINING FRICTION 


133 


the interface to the melting point, which would cor¬ 
respond to a P r of 3 to 10 klb/in. 2 during engraving. 
Further determinations of interface temperatures in 
conjunction with other ballistic and heat-input data 
might lead to valuable results in evaluating friction. 

6.2. 3 Direct Determination by 

Projectile Gauges 

Piezoelectric gauges 5 46 for the direct measurement 
of bore friction, acceleration, and base pressure and 
their methods of use are described in Section 4.5.3. 
Further mention should here be made of the bore 
friction gauge, which was essentially a pressure gauge 
mounted in the base of the projectile. An accelerated 
pressure gauge will measure not the pressure applied 
to the piston, but the difference between the pressure 
and the force per unit area needed to overcome the 
inertia of the piston. By constructing the piston so 
that the ratio of its mass to cross section equals that 
of the projectile, the force exerted by the piston on 
the crystal becomes proportional to the difference 
between the acceleration of the projectile and the 
pressure exerted on its base, that is, to the friction. 
Acceleration and base pressure gauges were similar, 
with effectively infinite and zero mass of the piston, 
respectively. 

The performance of the gauge did not equal the ex¬ 
cellent principle of its construction, mainly because 
of distortion of the signal. Some numerical results for 
the friction in a 20-mm Hispano-Suiza AA gun were 
obtained; the maximum P r for a number of rounds 
averaged 9.4 klb/in. 2 , and the friction quickly fell 
from this peak value attained in the first millisecond 
of travel to about 1.4 klb/in. 2 . These numbers are ex¬ 
tremely uncertain; the friction-time curve was not 
reproducible from round to round, and showed sev¬ 
eral sharp peaks, dropping nearly to zero, during the 
engraving period. If the difficulties of preventing 
spurious gauge vibrations and false signal pickup due 
to ionization could be overcome, these methods might 
be of considerable use. 

6 2,4 Semidirect Determination from 
Ballistic Firings 

The classical method of determining friction re¬ 
mains the best current one; that of analyzing ballistic 
data concerning the pressure of the powder gas and 
the position of the projectile. One prime objective of 
the ballistic firings 39,65,108 ’ 116 ’ 131 ’ 132 at Carderock of a 


Naval 3-in. gun, Mark VII, and of an experimental 
37-mm gun, T47, was the determination of the fric¬ 
tion. These firings are described in Chapter 4; we will 
here recapitulate the methods of obtaining and inter¬ 
preting the data relative to the friction. The subse¬ 
quent sections of this chapter present and discuss a 
few of the results. 

Determination of Projectile Acceleration 

The acceleration of the projectile was determined 
from either the position of the projectile, or the posi¬ 
tion, velocity, and acceleration of the gun in recoil. 
The recoil is proportional to the displacement of the 
projectile if the powder and powder gas move uni¬ 
formly and if there is no resistance due to the recoil 
brake while the projectile remains in the gun. 

Analysis of the recoil data, as obtained by the vari¬ 
ous instruments, leads to the conclusion that the dif¬ 
ferentiating instruments (rotary velocimeter, linear 
velocimeter, mutual inductance differentiator, crystal 
accelerometer) do not give results that may be ap¬ 
plied to the determination of friction, probably be¬ 
cause of vibrations in the gun or instruments. The 
step-by-step recoilmeter, however, gives results that 
are in excellent linear relationship to the best deter¬ 
mination of projectile position, after the first 6 in. of 
projectile travel, indicating that the recoil is essen¬ 
tially free and that the powder and gas are essentially 
uniformly distributed. 

The recoilmeter data, while of good accuracy, are 
not so extensive as the displacement data obtained 
by the microwave interferometer. Hence the latter 
are to be preferred in computing friction, when they 
have been obtained with good reliability, as in the 
3-in. gun; but the recoilmeter appears to be a simple 
and useful instrument, by which a number of points 
on the displacement-time curve may be obtained. 

The microwave interferometer gives essentially a 
continuous record of the projectile displacement, 
without modifying the projectile or gun in any signifi¬ 
cant way. In the 3-in. gun the projectile position was 
located at over 100 points per round from the start of 
travel to about 1 ft from the muzzle, with a relative 
accuracy of about 0.003 ft. The absolute accuracy as 
judged from the agreement of the microwave with the 
position given by barrel and ejection contacts may be 
no better than 0.02 ft; however the relative accuracy 
is the more significant in determining the accelera¬ 
tion. The time of arrival of the projectile at each point 
can be measured with an accuracy of 2 microseconds. 


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134 


BORE FRICTION 


From these data it should be possible to determine 
the acceleration with an accuracy of about 3 per cent 
(±10 ft/sec-msec). This corresponds to an uncer¬ 
tainty in P r of ± 0.5 klb/in. 2 . It is obviously neces¬ 
sary and desirable to reduce to a minimum the per¬ 
sonal factor and the accidental introduction of error 
in carrying out the double differentiation of the dis¬ 
placement-time data. Various methods of graphical, 
numerical, and analytical differentiation have been 
tested on the data for consistency and reliability. 
These are described in detail in the original reports 
and cannot be evaluated here. It is believed that least 
error is introduced when the displacement data are 
plotted as deviations from a “master curve” whose 
first and second time derivatives are known with ab¬ 
solute accuracy. Inasmuch as all similar rounds in a 
gun show nearly identical displacement-time curves, 
the deviations in displacement for any given round 
may be plotted on a greatly expanded scale, permit¬ 
ting increased accuracy in smoothing the experi¬ 
mental points and in differentiation of the smoothed 
curve. The smoothed curve may be drawn so that the 
velocity at the muzzle equals the velocity as deter¬ 
mined by measurements down the range. A correc¬ 
tion of 1 per cent or less, which may be attributed to 
the accelerating effect of the gaseous blast after the 
projectile has passed the muzzle contact, must some¬ 
times be applied in order to bring the internal and 
external velocities into best agreement. The smoothed 
curve may be located by least-squares methods, in re¬ 
gions where a simple analytical function may be used 
to represent the travel, as for example when the pres¬ 
sure-time, and hence the acceleration-time, curve is a 
linear or quadratic function. After the smoothed 
curve has been located on the deviation plot, its first 
and second time derivatives are found by graphical or 
numerical methods; these added to the known first 
and second derivatives of the master curve give the 
velocity and acceleration of the projectile. 


Determination of Base Pressure 

In the 3-in. gun, the gas pressure was measured 
simultaneously at four points along the barrel. The 
pressure on the base of the projectile is determinable 
by an extrapolation of the pressure-displacement 
curve from the gauge positions to the base. According 
to the Kent-Hirschfelder theory of gas flow and pres¬ 
sure distribution in guns referred to in Section 3.2.2, 
and according to a large number of similar theories to 
a degree of approximation small compared with the 


experimental uncertainty, the ratio of pressure at any 
point to the base pressure Px is given by equation (8), 

■^= 1 +2F( 1 -^ 

in which C is the weight of the charge, M is the weight 
of the projectile, and y is the ratio of volume between 
the point and the breech to the total volume between 
breech and projectile. By equation (8) we may thus 
calculate Px from the observed P at any gauge, and 
the known position of that gauge, as a function of the 
position of the projectile. When several gauges are 
recording simultaneously, the average calculated 
value of Px may be used. In case the observed pres¬ 
sures at the several gauges consistently disagree with 
equation (8), a similar relation, with an empirical 
constant r(y) replacing C/2M may be used. 

Having obtained the acceleration and the base 
pressure by the methods just sketched, the friction 
follows immediately by equation (1). Care must be 
taken to adjust the mass m of the projectile for the 
rotation of the projectile and the recoil of the gun, 
if the accelerations are computed relative to the 
gun. 


Smoothed Determination from 
Ballistic Measurements 


When applied to most existing sets of ballistic data, 
the procedures of Section 6.2.4 yield results for the 
friction which are quite erratic in time, and negative 
values,which are physically absurd, are often encount¬ 
ered. This is the natural result of having to determine, 
even from data of quite high accuracy, a small differ¬ 
ence between two much larger quantities, when there 
is some arbitrariness in determining one of these (the 
acceleration). In order to arrive at more reasonable 
values of the friction, it is frequently desirable to 
apply the following procedure, which has been termed 
the “integration method.” 

The “free velocity” F/, which is the velocity the 
projectile would acquire were there no retarding 
force, 

v,(t) = Pxdt ( 9 ) 

may be accurately evaluated from Px by equation 
(9). Then the corresponding “free travel” L f may be 
found from it by means of equation (10) 



V f {t)dt. 


( 10 ) 


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RESULTS—3-IN. GUN 


135 


The “retardation,” defined by equation (11), is 

X r = Lj — L (11) 

much smaller (from .05 — .25) than the observed 
travel L, but may be calculated with equal precision. 
By plotting against time not the observed travel but 
the retardation, we obtain both a smoothing of the 
travel data, by increase of scale, and the assurance 
that the friction will be everywhere positive, if we 
draw the curve through the observed points always 
concave upwards. Two differentiations of the smooth 
retardation-time curve give in turn the “retardation 
velocity” V r and the friction P r , in accordance with 
equations (12) and (13), respectively. 

V, ~W~ v '~ v ' ™ 


M_ dVr = p _ *L. 

gA dt x gA dt 


(13) 


That is, by subtracting out the base pressure before 
plotting and doubly differentiating the travel we ef¬ 
fectively smooth the data. Hence this method would 
appear to be the preferred means of determining fric¬ 
tion. 


6 2 6 Determination of Average Friction 

It is to be noted that two principal determining 
parameters of the retardation-time curve, namely, its 
value X rm and limiting slope V rm at ejection, do not 
depend on any internal travel data, but are given by 
the pressure-time curve, the known length of the gun, 
and the observed muzzle velocity. Hence two con¬ 
stants that characterize the average friction may be 
determined from these data alone, if a constant time 
relationship for the friction-time curve may be as¬ 
sumed. Similarly three constants characterizing the 
average friction may be calculated from the pressure¬ 
time curves, the muzzle velocity, and one internal 
travel point. 

For example, it was found 113 that a fairly good ap¬ 
proximation to the friction in the 3-in. gun was ob¬ 
tained by expressing it in terms of three constants, 
the engraving friction P re , the intermediate friction 
Ph, and the muzzle friction P rm . These are consistent 
functions of the reduced time t', the fraction t/to of 
the observed time to the total time to from start to 
ejection. The assumed functions are given in equa¬ 
tions (14). The three Pf s may be calculated from 
Xrm) Vrm and Xr at t — • 0.5. 


Pr = Pret'/0. 15 ( 0<*'<0.15)' 

= Pre ,(0.15 < f < 0.25) 

= Pre - (Pre ~ ~ 0.25)/0.15 


(0.25 < t' < 0.40) (14) 

= Pn -l(Pri - Prm)(t' ~ 0.4)/0.3 
(0.4 <t f < 0.7) 

= Prm (0.7 < ? < 1.0) 

A function similar to (14), with P ri = P rm is sug¬ 


gested as a suitable two-constant equation of general 
applicability, to calculate the friction from the pres¬ 
sure-time curve and the muzzle velocity. The equa¬ 


tions yielding the P re and P rm are then: 


Pr. = —Y-8.(>75 X 


Prm = 


—VJk ~ 2 ' 940 € 

MV V Y 

-~2.570^P - 3 . 2134 ^- 
qA to V m V m to 


( 15 ) 

(16) 


6 2 7 Correlation of Friction with Other 
Ballistic Measurements 

In Section 6.2.2 it was pointed out how the friction 
may be roughly inferred from measurements of strain 
or heat input. The more accurate methods may be 
checked against such measurements. For example, if 
friction has been determined by the methods of Sec¬ 
tions 6.2.4 or 6.2.5, from the results we may calculate 
the heating and melting of the rotating band; this in 
turn should correlate with observations on the band 
pressures and the coppering of the gun. The heat in¬ 
put due to friction may be calculated, and is to be 
considered in any theoretical accounting for the ob¬ 
served total heat input, as determined either by direct 
measurement of the temperature rise at various posi¬ 
tions along the gun, or by indirect calculation of the 
energy losses from the powder gas. 

If a series of projectiles of varying band diameter is 
fired, there should be a correlation with both the ob¬ 
served friction and band pressures. Any irregularity 
of the bore diameter as revealed by star-gauging 
should be reflected in the friction at that point. Thus 
in any comprehensive series of ballistic measurements 
the observation and correlation of these interlinked 
factors should be sought. 


6 3 RESULTS—3-IN. GUN 

6,3,1 Experimental Conditions 

The apparatus and methods by which were ob¬ 
tained the data necessary for the determination of 
friction in a 3-in. gun have been described in detail in 


CONFIDENTIAL 







136 


BORE FRICTION 


Chapter 4. The constants of the gun have been listed 
in Section 3.4.3. The rifling of the gun was of twist 
increasing from zero at the origin of rifling to one 
turn in 25 calibers at the muzzle. 

Of the 87 rounds fired from this gun at Carderock, 
approximately 35 were suitable for the determination 
of friction by the methods outlined in Sections 6.2.4 
and 6.2.5, in that experimentally reliable pressures 
were obtained at four positions along the gun, the 
position of the projectile was accurately located by 
the microwave interferometer, and the muzzle velo¬ 
city was known from the range solenoids. These rounds 
may be classified by the type of powder (NH, fired 
in the “Second Series” 65 from October 1943-May 
1944, or Pyro, fired in the “Third Series” 132 from 
October 1944-April 1945); by the fractional weight 


of service charge (1.0, 0.9, 0.8, 0.7, 0.6, and 0.5 for 
NH; 1.0, 0.75, and 0.5 for Pyro); by the condition of 
the bore surface (“greased” when the surface was cov¬ 
ered with a thin coating of heavy gun grease between 
firings; “dry” when the surface was untreated be¬ 
tween rounds, which were fired at intervals of from 
two hours to several days, with no “warm-up” 
rounds); or by the initial position of the projectile 
(“advanced” when it was uncrimped from the shell 
case so that the rotating band was rammed against 
the forcing cone; “normal” when it was left in place, 
giving it about 0.4 in. of run-up). Each change in 
these conditions altered to some extent the pressure- 
travel curves and the muzzle velocity, and hence the 
friction. Among rounds with identical conditions, 
variations also appeared. 



TIME FROM EJECTION IN MSEC 

Figure 2. Pressure-time curves for a greased round, full charge, second series in 3-in. gun. (NDRC Report A-323, Fig¬ 
ure 47.) 


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RESULTS—3-IN. GUN 


137 



cc 

z> 

ts> 15 





















ROUND 73 

GREASED 



























o L-ZJ ___________—------- 1 --=—^— 

13 12 11 10 9 8 7 6 5 4 3 2 1 0 -1 -2 -3 -4 -5 -6 -7 -8 -9 -10 

TIME TO EJECTION IN MSEC 


Figure 3. Pressure-time curves for greased and dry rounds, 50 per cent charge, third series in 3-in. gun. (NDRC Report 
A-460, Figure 36.) 



Figure 4 . Projectile displacement-time curves, microwave data, second series in 3-in. gun. (NDRC Report A-323, Fig¬ 
ure 32.) 


CONFIDENTIAL 







































































































138 


BORE FRICTION 


6 3 2 Typical Basic Data 

The basic data for the determination of friction are 
the pressures and the projectile displacement. The 
pressures were measured at hole 1, located in the 
chamber, 12.3 in. from the breech; at hole 2, 8.9 in. 
beyond the start of rifling or 116.9 in. from the muzzle; 
at hole 3, 75.3 in. from the muzzle; and at hole 4, 35.0 
in. from the muzzle. The pressures were read from the 
oscillograph records at intervals of 0.3 msec or less, 
and have been graphed as smooth pressure-time 
curves. A typical set of such curves, for a full-charge 
“dry” round of the third series, has already been pre¬ 



II 10 9876 543210 

DISPLACEMENT FROM MUZZLE IN FEET 

Figure 5. Pressure-travel curves for same round as Figure 2. (NDRC Report A-323, Figure 52.) 


sented as Figure 2 of Chapter 4; additional typical 
sets, for a full charge “greased” round of the second 
series, and 50 per cent dry and greased rounds of the 
third series, are exhibited in Figures 2 and 3. 

A reproduction of the oscillographic recording of 
the microwave interferometer for a 70 per cent d^ 
round of the second series, has been shown as Figure 
15 of Chapter 4. A plot of displacement from the muz¬ 
zle versus time, for dry 100 per cent, 70 per cent, and 
50 per cent rounds of the second series, is shown in 
Figure 4. From plots of the type illustrated the pres¬ 
sure and displacement may be read at evenly spaced 
time-intervals, and pressure-travel curves plotted, or 


CONFIDENTIAL 





































RESULTS—3-IN. GUN 


139 




Figure 6. Pressure-travel curves for same rounds as Figure 3. (NDRC Report A-460, Figure 39.) 


tables b of both functions constructed. The pressure- 
travel curves for the rounds illustrated by Figures 2 
and 3 are exhibited as Figures 5 and 6. 

As the first step in computing friction from the 

b Tables in which the pressure and displacement from the 
muzzle are given at intervals of 0.5 msec are given in NDRC 
Reports A-323 65 and A-460; 132 those reports also list the times 
of arrival of the projectile at the displacements corresponding 
to each of the 71 maxima and minima of the microwave, for 
several rounds. For some purposes, for example in applying the 
“integration” method for determining friction, it is more con¬ 
venient to have the pressure and travel from the starting posi¬ 
tion tabulated at a fixed number of even time-intervals. Such 
tables may be found in NDRC Report A-441, 113 at 20 values of 
the reduced time t' evenly spaced from the starting time to 
ejection. 


basic data, the pressure must be extrapolated from 
the observing gauge positions to the base of the pro¬ 
jectile. It is clear from a comparison of Figure 3 in 
Chapter 4 and Figure 5 of this chapter that there was 
a considerable difference in the magnitude of the 
pressure drop down the barrel between the two rounds 
depicted. This difference recurred consistently be¬ 
tween rounds of the second series, in which, for exam¬ 
ple, the average ratio Pi/ P 4 for all full-charge rounds 
when the projectile was 2 ft from the muzzle was 1.31, 
and rounds of the third series, in which the corre¬ 
sponding ratio was 1.14. The average pressure ratios 
for the two series are plotted against projectile travel, 
and compared with the simple theory [equation (8)] in 


CONFIDENTIAL 




































































140 


BORE FRICTION 



Figure 7. Pressure ratios versus travel, 3-in. gun, second and third series. (NDRC Report A-441, Figure 1.) 


Figure 7. The theory represents the results of the 
third series quite well, and is used for the extrapola¬ 
tion to the base of the projectile; for the second series 
an empirical extrapolation was required. Since no 
theoretical explanation of the wide discrepancy in the 
pressure-drop behavior in the two series has been 
offered, it is most probable that there is an experi¬ 
mental uncertainty in one set of data, although the 


error should not have exceeded 5 per cent or 1,000 psi. 
Tables of Px have been issued. 113 

6 3 3 Typical Results for Friction 

A determination of friction by the method which 
appears most reliable, namely the integration method 
of Section 6.2.5, is illustrated for a typical 100 per 


CONFIDENTIAL 




























RESULTS—3-IN. GUN 


141 


cent dry round of the third series in Figure 8. The re¬ 
tardation X r is plotted in two segments in order to 
permit a more expanded scale. The first segment runs 
from the start, at about 8 msec before ejection, to 4 
msec, when the free travel is 1.966 ft, the observed 
travel 1.633 ft, and hence the retardation 0.333 ft. 
The retardation velocity V r has increased from 0 to 
172 fps in this interval. From 4 msec to ejection V r 
further increases to 271 fps, corresponding to V m 
= 2,728 fps; in this time interval instead of X r we 
plot Xr = X r — 27 It. Each measured microwave 


point is plotted; these fall closely upon the smooth, 
concave-upward curves drawn to represent X r . The 
deviations in X r are less t^ian 0.005 ft, except for a 
few points nearest the muzzle. The curve, however, 
does not intersect the X r ' axis at t = 0; the deviation 
of 0.024 ft, equivalent to 9 msec, represents a typical 
experimental discrepancy between the location of the 
projectile as given by the microwave and by the muz¬ 
zle contact. The retardation velocity V r is obtained 
by graphical differentiation of X r , and is probably 
accurate to 3 fps, except near the muzzle, where an 



Figure 8. Determination of friction, integration method; 3-in. gun, third series, round 75, dry, full charge. (NDRC Re¬ 
port A-441, Figure 4.) 


CONFIDENTIAL 
































142 


BORE FRICTION 



0 0.2 0.4 0.6 0.8 1.0 

t'~ FRACTIONAL TIME TO EJECTION 


Figure 9. Simplified determination of friction, integration method; 3-in. gun, second series, round 44, dry, 70 per cent 
charge. (NDRC Report A-441, Figure 6.) 


uncertainty of 10 fps may occur. The curve is drawn 
to agree with the extrapolated range velocity, assum¬ 
ing no acceleration of the projectile after leaving the 
muzzle. Graphical differentiation of V r yields the 
friction P r ; the probable error here may not exceed 
5,000 ft/sec 2 in dV r /dt, or 300 psi in P r . However any 
uncertainty in P x must also be taken into account. 
Hence, the secondary rise in friction near 3 msec, al¬ 


though clearly required by the travel data as plotted, 
may not be real although it may be related to cop¬ 
pering (Section 10.5.4). The course of P r versus time 
is typical of the third series; it has already been ex¬ 
hibited, compared with the pressure, and plotted 
against the travel, in Figure 1. 

A typical friction determination for the second ser¬ 
ies (70 per cent dry) is presented in Figure 9. Here the 


CONFIDENTIAL 



















RESULTS-3-IN. GUNS 


143 


reduced time scale t' is used, and the X r curve is based 
not on the individual microwave points, but on 
smoothed values of the travel at 20 intervals of t'. In 
addition to the V r and P r curves obtained by graph¬ 
ical differentiations of X r , there are shown (dashed 
lines) the V r and P r resulting from application of the 
three-constant approximation equation (14). The 
simplified averaging method is seen to give a fairly 
good approximation to the detailed results. 

The energy of friction E f for the two rounds illus¬ 
trated, and the corresponding ratios c of E f to the 
total kinetic energy of projectile and powder, are 
given in Table 1. 


Table 1 . Typical frictional energies and ratios for 3-in. 
gun. 


Reduced time 

f — 

.25 

.5 

.75 

1.0 

Round 75 Ef (10 3 

ft-lb) — 

2.94 

26.5 

79.4 

107.5 

(Third Series) 

c — 

0.207 

0.062 

0.067 

0.065 

Round 44 F/(10 3 

ft-lb) — 

4.81 

22.5 

39.0 

42.6 

(Second Series) 

c — 

0.997 

0.119 

0.058 

0.042 


The heating of the rotating band in round 75 has 
been calculated from the friction by the methods 
sketched in Section 6.2.2. The result is that about 60 
per cent of E f is expended in heating the bore, 15 per 
cent in heating the band, and 25 per cent in melting 
the band, the melting occurring at travels from 5 to 
12 in. and 18 to 85 in. This is in general accord with 


the observed coppering of the gun, which is notice¬ 
able at hole 2 (9 in. travel) but much heavier at hole 3 
(50 in. travel). The calculated amount of melting in 
the second period appears to be too high, leading to 
the conclusion either that the experimental P r is too 
high (due to error in P x ), or that retarding forces 
other than band-bore friction, such as body engrav¬ 
ing, are operative. 

Other typical results, such as graphical depiction 
of friction-time and friction-travel data for other 
rounds, and tabulations of P r at 20 time-intervals 
and E f and c at four time-intervals for many rounds, 
may be found in the detailed reports. 65 * 113 ’ 132 

6 3 4 Summary and Discussion of Results 

The principal results of the firings of the 3-in. gun, 
as related to bore friction, are summarized in Table 2. 
This gives, averaged for rounds of each type fired, the 
muzzle velocity V m relative to the gun; the maximum 
pressure in the chamber P p ; the travel at maximum 
pressure L p ; the friction averaged over the engraving 
period, from 0.2 to 2 in. travel, P re ; the friction aver¬ 
aged over the intermediate range of travels from 2 to 
25 in., P ri ; the friction averaged over the rest of the 
bore to the muzzle, P rm ; the total energy of friction, 
Ef m ; and the final ratio of frictional to kinetic energy, 

Cm- 

From Table 2 the effect of changing the conditions 
of firing may be seen. The effect on engraving friction 


Table 2. Summary of results pertaining to bore friction from firings of 3-in. gun. 


Powder 

Charge 

Boref 

V m 

fps 

P P 

klb/in. 2 

Lp 

in. 

Pre 

klb/in. 2 

Pri 

klb/in. 2 

Prm 

klb/in. 2 

Ef m 

10 3 ft-lb 

Cm 

Pyro 

1.0 

G 

2698 

40.2 

17.2 

2.5 

1.81 

1.51 

121.7 

.0749 

Pyro 

1.0 

D 

2712 

41.6 

15.7 

2.9 

2.30 

1.41 

123.9 

.0757 

Pyro 

0.75 

D 

2269 

27.9 

14.2 

6.6 

2.18 

1.03 

97.6 

.0726 

Pyro 

0.5 

G 

1730 

15.0 

16.2 

3.3 

1.55 

0.68 

68.8 

.1144 

Pyro 

0.5 

D 

1867 

21.4 

9.8 

7.0 

2.29 

0.87 

91.7 

.1246 

NH 

1.0 

G 

2730 

43.6 

20.9 

1.9 

-0.1? 

0.30 

26.4 

.0160 

NH 

1.0 

D 

2748 

47.1 

19.1 

4.2 

0.6 

0.43 

45.3 

.0267 

NH 

1.0 

G* 

2713 

39.2 

26.2 

1.1? 

-0.5? 

0.1? 

0 ? 

.0 ? 

NH 

1.0 

D* 

2724 

42.0 

20.1 

0.3 

-0.4? 

0.03 

-17.5? 

-.0106? 

NH 

0.9 

G* 

2504 

31.1 

25.9 

1.0 

-0.1? 

0.32 

16.1 

.0117 

NH 

0.9 

D 

2571 

40.0 

16.0 

4.7 

1.6 

0.55 

47.9 

.0330 

NH 

0.8 

G 

2315 

27.0 

22.0 

2.1 

1.35 

0.94 

73.9 

.0634 

NH 

0.8 

D 

2387 

32.3 

18.2 

3.3 

0.7 

0.67 

57.9 

.0468 

NH 

0.7 

G 

2125 

21.4 

21.0 

2.5 

0.8 

0.33 

31.3 

.0322 

NH 

0.7 

D 

2150 

24.5 

20.8 

4.1 

1.55 

0.32 

42.6 

.0424 

NH 

0.6 

G 

1954 

18.1 

20.0 

2.3 

0.72 

0.46 

40.5 

.0498 

NH 

0.6 

D 

1986 

21.0 

17.3 

4.2 

1.78 

0.60 

63.9 

.0760 

NH 

0.5 

G 

1774 

16.1 

15.0 

2.7 

1.35 

0.52 

54.0 

.0813 

NH 

0.5 

D 

1842 

18.0 

13.2 

4.7 

1.38 

0.27 

33.6 

.0467 


* Projectile in normal position; 0.3-in. free run-up. 
fG = greased; D = dry 


CONFIDENTIAL 











144 


BORE FRICTION 


is particularly to be marked, this being the only fric¬ 
tional parameter determined with any great accur¬ 
acy. since the error may reach 1.0 klb in. 2 . In both the 
second and third series P re was lower for rounds with 
greased bore than with dry bore, and lower for those 
few rounds with run-up than when the projectile was 
initially advanced against the forcing cone. 

The effect of decreased velocity entering the forcing 
cone on increasing the engraving friction is further 
shown by the increase of P rf with decreasing charge. 
This effect is especially noticeable in some of the 50 
per cent dry rounds, where the friction exceeds the 
powder pressure at a travel of 0.3 to 0.5 in. to such an 
extent that the projectile decelerates and nearly 
comes to a stop. That the engraving friction has a 
profound influence on the subsequent ballistic course 
is demonstrated by the fact that when P rt is high 
there are higher maximum pressures, reached at 
shorter travels, and resulting in higher muzzle velo¬ 
cities. The same result is observed for individual 
rounds within a type; Table 3 shows the variation in 
the 75 per cent rounds of the third series. 


Table 3. Variation of results associated with engraving 
friction (Pre) 75 per cent dry rounds, third series, 3-in. 
gun. 


Round 

T m 

P, 

L, 


No. 

fps 

klb in. 1 

in. 

klb in. 5 

64 

2244 

25.5 

16.6 

3.8 

68 

2262 

28.0 

16.2 

5.1 

76 

2308 

30.2 

14.8 

6.4 


The intermediate and muzzle frictions P„ and Prm 
are less accurately known, and are of less ballistic- 
consequence, except that they are the principal con¬ 
tributors to the energy of friction. There does not ap¬ 
pear to be any marked correlation of the bore friction 
with the state of the bore surface; with decreasing 
charge the energy loss decreases slightly but the ratio 
of frictional to kinetic energy increases markedly. 
The most marked effect is the difference between the 
second and third series, but this is hardly likely to be 
a real effect, being most probably due to the same 
experimental uncertainty that gave rise to the unex¬ 
plained difference in the pressure drop. For this reason 
it is hard to say what the best average value of c for 
the gun may be; since the high frictions of the third 
series lead to implausibly large amounts of band 
melting, that series may be too high; but the second 
series bore frictions are likewise implausibly low for a 
gun with increasing twist. The most probable result. 


for full charge, probably lies between the two ex¬ 
tremes. namely c = .05 ± .03. 


« RESULTS. 37-MM GUN WITH 
PRE-ENGRAVED PROJECTILES 

* 4 * 1 Experimental Conditions 

The 37-mm gun. T47. is described in Sections 4.2.3 
and 31.7. The six lands for the pie-engraved projec¬ 
tile had a constant twist of 24 calibers per turn. The 
chamber volume was 31.43 cu in.: the travel of the 
projectile from seating to ejection. 73.96 in. Forty- 
three rounds were fired, at least three in each of 12 
classes, which differed in the type of charge and pro¬ 
jectile. as listed in Table 4. 


Table 4. Conditions of firing 37-mm gun, T47. 


Class 

Powder Primert 

Tvpe* Weight 
lb 

Projectile J 
Weight Length 
lb in. 


1 

Ml 

0.819 

>1 

1.62 

6.25 


2 

Ml 

0.819 

§ 

1.62 

6^5 

PL 

3 

Ml 

0.819 

M 

1.62 

5.75 


4 

Ml 

0.819 

L 

1.62 

5.75 


5 

Ml 

0.797 

5 

1.62 

6.44 

Obt 

6 

M5 



1.62 

6.25 


7 

M5 

0.828 

Js 

1.62 

6.25 

PL 

8 

M5 

0.828 

L 

1.62 

0.75 


9 

M5 

0.803 

s 

1.62 

6.44 

Obt 

10 

M5 

0.850 


1.34 

5.75 


11 

M5 

0.839 

s 

1.62 

6.25 


12 

M5 

0.794 

g 

1.92 

6.25 


# Some 

? powder properties were: 

T, 

F 

T 

W 




K 

ft-lb lb 


in. 



Ml 

2406 

-302.100 

1-239S 

-Q231 



M5 

3278 

363.380 

1-2284 

.OKS 


- - - s = s : i : - 

M = Medium. M38B2. lensti 28 m. 

L = T/gi; T3t leogtk 7.7 ia. 

jTwo special types ctf projerl&s mere iadoded: PL = Putfr-Lutoeid. 
that k surface treated wixfe iron pke^ote oaatrae: Obt = Obturated, 
with a i- ttitti copper shirt near the base. 

Pressure-time records were obtained by two gauges 
located at opposite ends of the same diameter in the 
chamber, at about four-fifths of the distance from the 
breech to the seated projectile. The gauges served as 
experimental checks on each other, but gave no in¬ 
formation concerning the pressure gradient down the 
barrel. It was therefore necessary to compute the 
base pressure by the theoretical equation (8!. The 
displacement, velocity, and acceleration of the pro¬ 
jectile were determined by the microwave interfero¬ 
meter : in this gun the records were at times disturbed, 
possibly due to ionized gases that had leaked past the 


CONFIDENTIAL 











RESULTS, 37-MM GUN WITH PRE-ENGRAVED PROJECTILES 


145 



projectile. An unexplained dependence of the type of 
microwave record cm the type of projectile also oc¬ 
curred. the results being most satisfactory for classes 
10 and 12. least for classes 1. 9. and 11. Typical pres¬ 
sure-time curves for Ml powder, showing how igni¬ 
tion waves are reduced by increasing the primer 
length, are presented in Figure 10. Typical micro- 
wave records of five classes are shown in Figure 11. 

tu Results and Discussion 

Because of the uncertainty in base pressure (of the 
order of ±2 klb in.-) imposed by the use of only one 
gauge position and the experimental error at that 
point, and in projectile displacement and acceleration 
caused by the erratic microwave records, the results 


of friction determinations in the 37-mm gun, T47 are 
not of high accuracy. The results for a typical round 
of class 12, as determined by the integration method 
(Section 6.2.5), are shown in Figure 12. The individual 
microwave points scatter rather widely about the 
smooth curve drawn to represent Xr, the average 
deviation being 0.008 ft. In view of the shorter time 
interv als in this gun, this would lead to a probable 
error of about ± 15 fps in V T , and + 30,000 ft/sec 2 
in dV r dt or ± 1 klb/in. 2 in P T . For classes of rounds 
in which the microwave record was less satisfactory, 
the scatter and. resulting inaccuracy are greater. 
There is, however, one consistent feature of the re¬ 
sults which is illustrated in this typical round, 
namely, that the friction is low* at the start, which 
is the reason that pre-engraved projectiles reduce 


CONFIDENTIAL 























































146 


BORE FRICTION 



/ / 

/ 3 

( i ^ '' ❖ »' it i * 

/\ 

A * • * * * * • 

-- -- — ^ V 

J * . ♦ < . . 

n 





■v 




EXCELLENT 



t 3 r ? 

/ v v, / A Z' '• 

.-^ \ / VV . 

( # 0 0 " 

■ ' ■ ? ' . 

* /» 

V 

S & £ X 

>. . VP ^ 

♦ ? 


•s . * » 

* * * « ‘ <<<(*'’ 


VERY GOOD 



Figure 11. Microwave records for five classes of projectiles in 37-mm gun: Excellent—Class 10; very good—Class 4; 
good—Class 2; fair—Class 5; erratic—Class 9. (This appears as Figure 4 in NDRC Report A-459.) 


CONFIDENTIAL 










RESULTS, 37-MM GUN WITH PR-ENGRAVED PROJECTILES 


147 



erosion (Chapter 31). The friction reaches a maximum 
near the time of maximum acceleration or later. 
The difference from the results in the 3-in. gun, due 
to the absence of the rotating band, is marked. 

A summary of the results is presented in Table 5. 
This gives, averaged for each class, the time from 
start of pressure to ejection, the muzzle velocity, the 
maximum pressure, the friction averaged at 0.5 msec 
intervals, and the final ratio of frictional to kinetic 
energy. 

Within the limited accuracy of the results, no def¬ 
inite correlation can be made between the friction and 
the type of projectile. Very little change in the over¬ 
all ballistics is caused by “Parco-Lubrizing” the pro¬ 
jectile surface, or by adding the thin obturating ring. 
The comparatively high resisting pressures found with 
all classes can hardly be due entirely to projectile- 
bore friction. The observed heat transfers were not 
anomalously high, nor did they show a maximum near 


the point of maximum acceleration, indicating that 
the large apparent P r did not give rise to a corre¬ 
spondingly large heat of friction. The high values 


Table 5. Summary of results pertaining to friction, 
37-mm gun. 


Class* 

to 

msec 

V m 

fps 

P P 

klb/in. 2 

Pr, (av.) 
klb/in. 2 

C m 

1 

3.86 

3348 

57.8 

1.7 

.071 

2 

3.92 

3347 

58.7 

2.1 

.070 

3 

3.96 

3328 

57.4 

2.5 

.087 

4 

4.48 

3261 

52.0 

2.4 

.087 

5 

3.96 

3290 

55.3 

2.8 

.098 

6 

4.15 

3588 

63.1 

2.5 

.092 

7 

3.98 

3595 

63.4 

3.4 

.099 

8 

4.07 

3604 

63.5 

2.9 

.074 

9 

4.04 

3513 

56.6 

1.7 

.050 

10 

3.74 

3868 

59.2 

1.7 

.063 

11 

3.90 

3631 

64.1 

2.1 

.081 

12 

4.33 

3302 

64.3 

3.1 

.096 


* See Table 4 for a definition of each class of rounds. 


CONFIDENTIAL 
























148 


BORE FRICTION 


found may be due in part to errors in measurement 
and calculation of Px ; or in part it may represent a 
resisting pressure on the part of the projectile caused 
by the leakage of gases. Several lines of evidence, 
notably photographs of the muzzle before the pro¬ 
jectile has emerged, point to a considerable quantity 
of such gas. 

«•« DISTRIBUTION OF ENERGY OF THE 
POWDER 

651 Methods of Determining 

Energy Balance 

The Energy Equation 

Another important result of a complete set of bal¬ 
listic measurements, such as those carried out at Car- 
derock, is that several lines of experimental evidence 
are brought to bear on the problem of the energy dis¬ 
tribution of the powder. A theoretical discussion of 
the ballistic energy equation is given in Section 3.2.3; 
for the present purposes we may write the funda¬ 
mental equation (17), which is essentially the same as 
equation (6) of Chapter 3, 

E - - T^iO - t) ~ E - 

M C 

= 2^ + & V2 + E ' +h - W 

This equation expresses the fact that the extent by 
which the gas temperature T falls below the flame 
temperature T 0 is a measure of the energy released as 
work or heat, or unexpended due to incomplete burn¬ 
ing; the terms on the second line distribute the re¬ 
leased energy among the kinetic energy of the pro¬ 
jectile, the kinetic energy of the powder and gas, the 
energy of overcoming friction, and the heat trans¬ 
ferred to the walls. In the Carderock firings all the 
terms of equation (17) were determined, directly or 
indirectly; a more detailed analysis than the brief 
summary to be presented here may be found in the 
original reports, especially A-441 113 and A-444. 116 

The powder constants F, y , and T 0 have been cal¬ 
culated by both the detailed and approximate (addi¬ 
tive) methods sketched in Section 2.4. The inaccuracy 
introduced by use of the simpler approximate method 
is less than 3 per cent, except at times near shot ejec¬ 
tion, when the temperature is below 1750 K and sec¬ 
ondary gas reactions leading to methane formation 
become of importance. 


The kinetic energy terms, which constitute the 
largest fraction of the expended energy, may readily 
be calculated from the projectile velocity, which was 
determined as a function of time by the methods de¬ 
scribed in Section 6.2. The factor 8 may be taken as 3, 
the theoretical value, with sufficient accuracy, except 
at times near the start of travel when the turbulent 
velocity of the unburned powder considerably ex¬ 
ceeds the projectile velocity. E f may also be calcu¬ 
lated by the methods described in Section 6.2. 

The problem of determining energy balance thus 
depends on a knowledge of N, T, E u , and h. At times 
near shot ejection, the powder is completely burned, 
and A r = C. At such times, with rare exceptions to be 
mentioned in the next section, E u may be neglected. 
Hence the problem is solved if either T or h may be 
determined or calculated. 

Average Temperature 

The average temperature T of the gas throughout 
the gun volume may be estimated from the experi¬ 
mentally determined temperatures (Sections 2.5 and 
4.3.17), but the results are of limited accuracy, except 
at position 1, in the powder chamber. Even there, be¬ 
cause of the high degree of turbulence and tempera¬ 
ture inhomogeneity in the gas, values averaged over 
a number of rounds are needed to obtain representa¬ 
tive results. The observed temperatures at the for¬ 
ward positions appear to be considerably lower than 
the average temperature at those points, due to cool¬ 
ing by the walls. 

The average temperature T may be approximated 
roughly by the chamber temperature T\ and also it 
may be calculated from the observed average pres¬ 
sure [computed from the observed pressures by the 
use of equation (8) or other empirical representation 
of the pressure gradient] and travel, by means of the 
equation of state, equation (15) of Chapter 3. The ex¬ 
perimental value of Ti (or of h) may be used to check 
the covolume rj ; in general it is better to assume that 
7] is accurately calculable by the methods of Section 
2.4, and to use the equation of state to check the ex¬ 
perimental self-consistency of P, Ti, and h, or to cal¬ 
culate T and h from P. 

Heat Input (h) 

The heat h transferred from gas to bore may be cal¬ 
culated by the methods described in Section 5.3. The 
computation may either follow the detailed method 


CONFIDENTIAL 




DISTRIBUTION OF ENERGY OF THE POWDER 


149 


based on the fundamental equation (18), 

^ = i\( T, - T.)C p Av, (18) 

using an assumed value for the friction factor X, and 
the observed ballistic data for the gas temperature 
T 0 , density A and velocity v, or use may be made of 
tables 48 based on the observed muzzle velocity and an 
assumed simple ballistics. It has been shown 116 that 
the two methods of calculation of h were in good 
agreement for the 3-in. gun, and that h was very 
nearly a linear function of the travel L. 

The quantity h may also be derived from experi¬ 
mental determinations of heat input at various points 
along the gun, as described in Section 5.4. This also 
requires theoretical calculations concerning the con¬ 
tributions to the observed total heat input of the heat 
of friction and of the heating by the gases after ejec¬ 
tion. Using either ballistic data or theoretical compu¬ 
tations regarding the state of the powder gas after 
ejection, the result is obtained that the heat input up 
to ejection, h m , is approximately 0.64 of the total heat 
input. For times before ejection, a linear dependence 
of h on L is assumed; the alternative assumption, 
equation (10) of Chapter 3, made in the theoretical 
ballistics, may also be used. 

Energy Balance During Burning 

At times before the powder is completely burned, 
any two of the three quantities P, T , and h must be 
known in order to calculate N/C from equations (18) 
in this chapter and (15) in Chapter 3, if E u is assumed 
negligible; or all three must be known to calculate 
E u . In general the best procedure is to make use of P 
and h, assuming E u = 0; if T calculated in this way 
is within 100 Kelvin degrees of the observed T, the 
energy balance is reasonably reliable. If the observed 
T falls below the calculated by more than 100 de¬ 
grees, it must be associated with an E u > 0. 


6 5 2 Energy Balance during Firing 
of 3-inch Gun 

At Ejection 

To illustrate the results of the different possible ex¬ 
perimental and theoretical approaches to the compu¬ 
tation of the energy balance at the time of ejection, 
there are presented in Table 6, for various conditions 
of firing, values of T calculated by the following 
methods. 

1. The experimentally determined temperature in 
the chamber; 

2. Values calculated from the experimentally de¬ 
termined average pressure, using alternatively: 

a. The approximate additive equation of state 
(15) in Chapter 3, 

b. The more exact equation (10) of Chapter 2; 

3. Values calculated from the energy balance equa¬ 
tion (17) of this chapter, assuming E u = 0, and using 
alternatively: 

• 

a. The approximate method of calculation of 
E r ei and the observed value of h, 

b. The exact calculation of E reh and the observed 
value of h, 

c. The exact E re i and the value of h calculated 
using the simplified ballistics given in Section 
3.2.7 and the value of X corresponding to 
Q = 10.2 cal/cm 2 in the caliber .50 gun. 

From Table 6, supported by the more detailed com¬ 
putations not reproduced here, the following conclu¬ 
sions may be drawn. 

1. The various experimental results do not lead to 
identical values of T, but the disagreement is not 
more than ± 5 per cent, and hence is attributable to 
experimental error in the determination of tempera¬ 
ture and pressure. 

2. The differences between calculations based on 


Table 6. Average values of gas temperature at shot ejection, 3-in. gun. 


Powder— 

Charge— 

Bore—(G = greased; D = dry) 
Source and method of calculation of T, (K) 

NH 

1.0 

G-D 

NH 

0.7 

D 

NH 

0.5 

D 

Pyro 

1.0 

G-D 

Pyro 

0.75 

D 

Pyro 

0.5 

D 

Pyro 

0.5 

G 

1. Experimental 

1634 

1740 

1518 

1696 

1752 

1650 

1590 

2a. Pressure, approx eqn of state 

1689 

1682 

1575 

1721 

1686 

1642 

1675 

2b. Pressure, exact eqn of state 

1706 

1690 

1579 

1737 

1692 

1646 

1678 

3a. Energy, approx calc; h obs. 

1610 

1683 

1639 

1597 

1640 

1680 

1732 

3b. Energy, exact calc; h obs. 

1651 

1707 

1653 

1626 

1657 

1687 

1737 

3c. Energy, exact calc; h theor. 

1629 

1677 

1611 

1603 

1624 

1642 

1697 


CONFIDENTIAL 







150 


BORE FRICTION 



Figure 13. Energy distribution of powder, 100 per 

cent charge, 3-in. gun. 

the approximate and detailed computation of covol¬ 
ume and energy release are less than the differences 
between the several experimental methods; hence 
more exact measurements are needed before an ex¬ 
perimental determination of the covolume, or other 
checks on the finer points of the calculations can be 
made. 

3. For all rounds except at 50 per cent charge, bet¬ 
ter agreement with the other temperatures is ob¬ 
tained by using the experimental rather than the 
theoretical values of h ; that is, the other results rein¬ 
force the experimental heat inputs and suggest that 
in this gun a value of X from 20 to 25 per cent lower 
than that found in the caliber .50 gun was operative. 

4. For the 50 per cent rounds, and most markedly 
for the greased Pyro rounds that gave the lowest P p 
and V m (Table 2), the energy balance gives tempera¬ 
tures that exceed the experimental temperatures by 
amounts greater than the probable error. Hence in 
these cases there is an energy deficit E u to be attrib¬ 
uted to incomplete reaction of the powder. In such 
rounds an E u of similar magnitude is also found at 
times before ejection, by methods stated in Section 
6.5.1. 

The energy distribution at ejection is summarized 
in Table 7. The various energies are expressed as per¬ 
centages, first of the total potential energy the pow¬ 
der could liberate if cooled to 0 K, and second of the 


kinetic energy of the projectile. The total heat trans¬ 
mitted to the gun, which includes in addition to h m 
the contributions of frictional and post-ejection heat¬ 
ing, is also listed. Figure 13 illustrates the overall 
result. 


Before Ejection 

At times before ejection, the energy distribution is 
roughly similar to that just discussed. Some values of 
the ratio c of frictional to kinetic energy have been 
presented in Table 1. The ratio 0 of heat loss to kin¬ 
etic energy [see equation (11) of Chapter 3 in Section 
3.2.3] in general shows a minimum at times near 
t f = 0.6. Some typical results are presented in Table 
8. This gives, for the average full-charge firing of NH 

Table 8. Observed and calculated gas temperatures— 
100 per cent charge NH powder, 3-in. gun. 


Values of T, K 


f 

Exp. 

Calc. 

h 

Calc. 

7 = 1.275 

Calc. 

7 = 1.30 

N/C 

0.05 

2045 

2636 

2636 

2636 

0.0278 

0.10 

1948 

2635 

2636 

2636 

0.0467 

0.15 

2267 

2631 

2634 

2634 

0.0969 

0.2 

2520 

2621 

2625 

2624 

0.1650 

0.3 

2528 

2569 

2574 

2569 

0.3120 

0.4 

2413 

2448 

2451 

2436 

0.5358 

0.5 

2373 

2296 

2295 

2268 

0.7691 

0.6 

2192 

2156 

2155 

2118 

0.9700 

0.7 

2010 

1984 

1988 

1934 

0.9882 

0.8 

1869 

1839 

1856 

1785 

1.0033 

0.9 

1760 

1716 

1749 

1768 

1.0004 

1.0 

1634 

1605 

1662 

1573 

1.0157 


powder, at various times, the temperature as observed, 
and as calculated from the observed pressure, travel, 
and heat loss using the experimental value of h and 
two different assumptions for a constant ratio 0. The 


Table 7. Distribution of energy of the powder, 3-in. gun. 


100 Per cent charge 50 Per cent charge 


Type of energy 

E 

10 3 ft-lb 

Per cent 
of total 

Per cent 
of {KE) P 

E 

10 3 ft-lb 

Per cent 
of total 

Per cent 
of (KE) i 

Total potential of powder 

CF /{7 - 1) 

5144 

100.0 


2572 

100.0 


Released at ejection 

CF(1 - T m /T 0 )/(y - 1) 

1984 

± 100 

38.6 

135.1 

1016 ± 100 

39.5 

160.0 

Kinetic of projectile 

(KE) P 

1469 

± 20 

28.6 

100.0 

635 ± 60 

24.7 

100.0 

Kinetic of powder gas 

C KE)g 

148 

± 20 

2.87 

10.1 

32 ± 5 

1.24 

5.0 

Kinetic of recoil 

(KE) r 


11 

0.21 

0.7 

4.0 

0.16 

0.6 

Rotation of projectile 

Erot 


12 

0.23 

0.8 

5.3 

0.21 

0.8 

Frictional 

E f 

80 

± 30 

1.56 

5.4 

65 ± 20 

2.53 

10.2 

Gaseous heat transfer 

h 

264 

± 100 

5.13 

18.0 

181 ± 80 

7.04 

28.5 

Unreleased potential 

E u 


0 

0 . 

0 . 

94 ± 90 

3.65 

14.8 

Total heat to gun 

H 

482 

± 50 

9.37 

32.8 

338 ± 40 

13.14 

53.2 


CONFIDENTIAL 

























CONCLUSIONS 


151 


fraction burned as calculated in this way is also listed. 
It is seen that the usual ballistic assumption of con¬ 
stant 0 does not lead to wide errors in the calculated 
temperature; the experimental results for both T and 
h favor a lower ratio than was usually adopted 
for ballistic calculation, as suggested in Table 2 of 
Chapter 3. 

A noteworthy feature of Table 8 is the large extent 
by which the experimental T falls below the theoreti¬ 
cal at early times. This has been attributed partly to 
an incomplete burning term E u , and partly to an in¬ 
crease in the kinetic energy of the solid powder mov¬ 
ing turbulently in the chamber. This temporarily lost 
energy is recovered by the gas as burning proceeds. 

The values of N/ C obtained in this way were used 
in the calculation of burning rates, as discussed in 
Section 2.2. 

6 5 3 Approximate Results for 

3 7-mm Gun, T47 

In the 37-mm gun, T47, the gas temperature was 
not measured, and so there is one less check on the 
consistency of the results. At ejection the tempera¬ 
tures calculated from the pressure averaged 9 per cent 
above those calculated from the energy balance with 
observed values of h. This adds additional weight to 
the possibility that the experimental pressures, and 
the friction calculated therefrom, were too high. 
Hence no very reliable results for the energy distribu¬ 
tion are at hand. The average value of the effective 
specific heat ratio y was 1.317 for Ml powder and 
1.289 for M5 powder, corresponding to 0 = 0.24, 
again somewhat lower than suggested in Table 2 
of Chapter 3. 

66 CONCLUSIONS 

The general conclusions regarding bore friction and 


related problems, that may be drawn from the experi¬ 
mental results and theoretical discussion of this chap¬ 
ter, may be summarized briefly. It appears that even 
with the data at hand, which in precision and com¬ 
pleteness compared very favorably with any previous 
ballistic results, it is difficult to obtain reliable results 
for such a small differential effect as the friction. It 
has however been demonstrated that the engraving 
friction in a 3-in. gun varies with the conditions of 
firing from 2 to 10 klb/in. 2 , or from 5 to 50 per cent of 
the maximum pressure, the variations in some cases 
being without assignable cause, and that these varia¬ 
tions are reflected in the maximum pressure and muz¬ 
zle velocity. 

The engraving friction bears a relation to the para¬ 
meter P 0 , the starting pressure, that appears in in¬ 
ferior ballistic calculations, but is not identical with 
it, inasmuch as the theoretical quantity is influenced 
by the various simplifying assumptions that are made 
concerning friction and burning rate. (See Section 
3.2.2.) The coefficient of bore friction c in this gun 
varies with the travel and the firing conditions, but in 
general appears to exceed the average value of 0.04 
assumed in the Division 1 theoretical ballistics. The 
energy distribution of the powder likewise varies with 
the travel; at early times and in low-density firings 
there is some additional energy deficit due to incom¬ 
plete evolution of powder energy. Otherwise the rela¬ 
tions are fairly in accord with theory, with the heat 
loss being somewhat lower than usually assumed. 

Further work of the kind described in this and the 
preceding chapters, using a variety of guns firing con¬ 
ventional and experimental projectiles and powders, 
is needed. Only in this way will it be possible to 
establish the general laws relating bore friction and 
powder energy distribution to interior ballistic theory 
and to gun design and performance. 


CONFIDENTIAL 



Chapter 7 

BAND PRESSURE AND RELATED STRESSES 
By H. L. Black a 


71 THEORY OF POWDER GAS AND BAND 
STRESSES IN GUNS 

711 Introduction 

D uring world war ii it became clear from actual 
failures of guns and shells during testing that the 
effect of rotating band pressure had not been given 
proper consideration in the design of either guns or 
shells. In addition, there developed a tendency to fire 
guns with powder pressures exceeding those for which 
they were originally intended. As a result a great deal 
of attention has been given to the problem of the 
stresses in a gun tube under combined band and pow¬ 
der pressures and the corresponding stresses in the 
wall of a shell. 

In this country this work has been carried on by 
the Army in the Office of the Chief of Ordnance 307 - 308 
at Watertown Arsenal, 249-256 for the Navy at 
Dahlgren Proving Ground 330 and at the Massachu¬ 
setts Institute of Technology, 326 - 327 and for Division 1, 
NDRC, at both the Catholic University of America 117 
and the National Bureau of Standards. 58 - 65 - 132 In Brit¬ 
ain an earlier interest in the subject was contin¬ 
ued 362,366,370,376,377,378,424 

As a result of the theoretical and experimental 
work discussed in the present chapter, important 
progress has been made in the determination of band 
pressure and of its effects in combination with pow¬ 
der pressure and thermal stresses. The knowledge 
gained will be a valuable aid to the designer of hyper¬ 
velocity guns. If muzzle velocities are to be pushed to 
levels formerly considered impossible, there is need of 
further study, particularly of the cause of pressure 
waves that may impose dangerously high sporadic 
stresses on the tube. 

The application of these results to improvements in 
the design of projectiles is taken up in Chapter 27. 
Some attention is given there to the first efforts to re¬ 
duce band pressure, a change that would appear to 
be a prerequisite to the use of the higher powder pres¬ 
sures that are required in hypervelocity guns. 


a Technical Aide, Division 1, NDRC. (Present address: De¬ 
partment of Mathematics, Michigan State College, East 
Lansing, Michigan.) 


712 Fundamental Assumptions 

In considering stresses in a gun tube, it is customary 
to make certain assumptions. 

1. The gun is a smooth-walled, hollow cylinder. 

2. The tube is sufficiently long (or the projectile 
base is considered at a sufficient distance from an end 
of the tube) to make end effects negligible. 

3. Band and powder pressures give axially sym¬ 
metric radial stresses. 

4. Under uniform radial pressures, the stress dis¬ 
tribution follows the equations of Lame, given in 
equations (1) to (3), 

Radial stress: rr = r ^ 1 4r~ M~~) (1) 

1 — /r \ o r r / 

Hoop stress »<: = | ^ + y) (2) 

in which E is Young’s modulus, n the Poisson ratio, w 
the radial deformation, r the distance from the axis of 
the tube, and l a constant of integration determined 
by boundary conditions. 

5. The generally accepted criterion for failure is 
that of Mises-Hencky, according to which the equiv¬ 
alent stress, S e , can be expressed by equation (4). If 
S e exceeds the tensile strength of the gun steel, failure 
of the gun is to be expected. 

2 S e 2 = (00 — rr) 2 + (rr — xx) 2 + (xx — 00) 2 (4) 

6. Stresses are algebraically additive. If, for exam¬ 
ple, the simultaneous stresses caused by powder pres¬ 
sure and by band pressure at a given point are known, 
the combined stress is given by their sum. In partic¬ 
ular, if the two stresses are graphed against position 
in the gun, the resultant stress may be obtained by 
composition of ordinates. This principle of superposi¬ 
tion can be extended to any number of stresses and is 
a valuable tool in developing the theory of stress dis¬ 
tribution. 


152 


CONFIDENTIAL 





THEORY OF STRESSES IN GUNS 


153 



MUZZLE END X s Z/T BREECH END 


Figure 1. Tangential strain in 3-in. gun tube as a function of distance from the band. (Figure 13, NDRC Report A-298.) 


7,1,3 Integrated Stresses in the Tube 

During the engraving of the rotating band of the 
projectile by the rifling of the gun, pressure is highly 
concentrated on a narrow strip, in width actually less 
than the band. However, stresses exist over the 
whole length of the tube, tending toward zero at the 
ends. Lame’s formulas for the deformation and the 
stresses resulting from uniform pressure are valid in 
this case also, if the uniform deformation is replaced 
by the integral of the deformation over the length of 
the tube and the uniform stresses by the integrals of 
the stresses taken over the length of the tube. 10 The 
resulting formulas are the analogues of equation (1), 
(2), and (3), with w, rr, and 66 replaced respectively 
by W, R, and 0 where W = J'wdx, R = J'rr e dx, 
and 0 = f 66dx. h 

An important principle employed in the develop¬ 
ment of this theorem is that every axially symmetric 
load distribution can be built up by superposition of 
half-infinite uniform radial loads. 

During engraving of the projectile the longitudinal 
stress in the gun tube is practically that for a closed 
chamber, but later it is less, and as the projectile 
nears the muzzle, the more nearly the stress problem 
approaches that of the open tube case. Formulas have 
been developed 13 for the axial stress at points between 
the projectile and the muzzle, between the projectile 
and the breech, and above the powder chamber 
while the projectile is moving down the bore. 


b Somewhat similar theorems have been presented by other 
writers. 512 ’ 657 


7,1,4 Methods of Determining Frictional 
Force from Band Pressure 


Three methods have been suggested 13 for determin¬ 
ing the frictional force 0 between a projectile and the 
gun barrel from a measurement of the strains in the 
barrel caused by band and powder pressure. The 
formulas used to determine the frictional force are 
based on the elementary solution of the stress equa¬ 
tions for an infinite, circular, cylindrical shell. The 
gun tube is acted on by both the uniform powder 
pressure P and a uniform longitudinal traction of 
total amount F. 

The analysis for one of the methods was extended 58 
and calculations were made to see whether it would 
be applicable to the 3-in. gun at Carderock, dsscribed 
in Section 4.2. The strain components on the outer 
surface of the tube may be represented by equa¬ 
tions (5) for gauge positions behind the projectile. 


W(z) 


>]e{ 2P 


rr 


m [F + M(z)F] \ 

Trri 2 y 

F + M{z)Y^ (5) 


— 2 M P + 


[r 2 2 (V) — ri 2 ] E ^ 7i tv 

Here, n and r 2 (z) are the inner and outer radii of 
the gun tube, M(z) is the mass of the tube between 
any plane z and the muzzle, Y is the positive accelera¬ 
tion of recoil of the gun, E is Young’s modulus, and ju 
is the Poisson ratio. These two equations are the 
equivalent of equation (6), which gives the friction, 
Ett [r 2 2 (z) — ri 2 " 1 


F = 


1 - 


+ e z »] — M(z)Y, (6) 


°This subject is dealt with more broadly in Section 6.2. 


CONFIDENTIAL 























154 


BAND PRESSURE AND RELATED STRESSES 



10 5 0 “5 -10 

MUZZLE END X = Z/T BREECH END 

Figure 2. Axial strain in 3-in. gun tube as a function of distance from the band. (Figure 14, NDRC Report A-298.) 


and equation (7), 

2 D Ett [r 2 2 (z) — rpj r , 

* Tl P = — 2(1 - m 2 )— 600 + ^ 7 

which gives the pressure on the base of the projectile 
directly from two measurements of strain behind the 
projectile. 

Equations (5) may be replaced by similar expres¬ 
sions in which the acceleration of recoil is expressed in 
terms of a “gauge constant.” This gauge constant was 
evaluated empirically from the firings of the 3-in. gun 
at Carderock, and then the curves shown in Figures 1 
and 2 were drawn for the tangential strain and axial 
strain, respectively, at a distance of 30 in. from the 
muzzle. d It was found that the curve for the tangen¬ 
tial strain would be unchanged if the friction were 
assumed to be zero and that even in the curve for 
axial strain the contribution from the friction is not 
large enough to be determined accurately. 

Therefore, it was concluded that friction cannot be 
determined directly from strain gauge measurements 
by this method. It was pointed out, however, that the 
use of strain gauges as pressure gauges should be 

d It was pointed out that, allowing for certain differences in 
the experimental conditions, these curves agreed well with 
some obtained 195 in the firing of a 76-mm gun, M1E2 at Aber¬ 
deen Proving Ground. 


satisfactory, and that acceleration of recoil might be 
evaluated in this manner. 

The other two methods 13 for determining friction 
from band pressure have not been similarly analyzed 
in terms of experimental conditions. The one involves 
the use of two longitudinal strain gauges set close to¬ 
gether over the powder chamber in a gun having a 
uniform w r all thickness in this region. The other 
method uses one such strain gauge and a recoilmeter. 

72 THERMAL STRESSES IN THE BARREL 
WALL 

7,2,1 Stresses at a Point 

If a nonuniform temperature distribution results 
during the heating or cooling of a solid body, the 
cooler portions of the body exert a restraint upon the 
expansive tendencies of the warmer portions and the 
mutual interference causes thermal stress which may 
be superposed upon any existing stress system within 
the body. This theory has been applied 151 in the case 
of a long, hollow cylinder unrestrained at the ends. It 
is assumed that plane sections far from the ends re¬ 
main plane and that the temperature field is axially 
symmetric and uniform along the axial coordinate x. 
The coefficient of thermal expansion a, the Poisson 


CONFIDENTIAL 





















THERMAL STRESSES 


155 


Table 1. Compressions at interface between a 14-in. stellite liner and steel caliber .50 machine gun barrel. 151 Case I, 
100-round burst; Case II, 250-round burst. Each case postulated for two situations: (a) moderate bore surface temper¬ 
ature; (b) very high bore surface temperature. 


1 


Compression 

Case I (a) 

(10 3 psi) 

Case I (b) 

(10 3 psi) 

\ 

Case II (a) 

(10 3 psi) 

Case II (b) 
(10 3 psi) 

Radial stress 

15 

20 

24 

26 

Maximum shear 

15 

12 

21 

20 

Tangential stress 

45 

47 

67 

62 


ratio (i, and the modulus of elasticity E are assumed 
invariant with temperature. With the boundary con¬ 
dition <j r — 0, at r = a and at r = b, where a and b 
are respectively the inner and outer radii, equations 
(8) and (9) are obtained for the stresses at a point r 
units from the axis. 

ar= j^~[-hf a aTrdr 

+ ?lb* - a*)L aTrdr i (8) 

a Trdr 

1 - M L r 2 Ja 

r 2 i a 2 f b “I 

. + <* Trdr ~ «t\. (9) 

Here T is the temperature excess over the lowest 
value in the body. 

7 2 2 Application to 

Caliber .50 Machine Gun 

Generalized stress equations were developed for 
composite hollow cylinders and the results were ap¬ 
plied to the computation of stress in a stellite-lined 
caliber .50 machine gun barrel (Chapter 22). Two 
temperature distributions based on experimental fir¬ 
ings (of the sort described in Section 5.5.1), were used; 
Case I, resulting from a 100-round burst, and Case II 
from a burst of 250 rounds. For each case two situa¬ 
tions, (a) and (b), were postulated within the liner; 
distribution (b) assumes a higher bore surface tem¬ 
perature and a steeper gradient than (a). Under the 
four sets of conditions, the thermal stresses in a gun 
with a Vi6-in. liner were exhibited graphically for three 
values of the Poisson ratio in the stellite. Case I (a) 
and Case I (b) were also graphed for a liner of thick¬ 
ness V 32 -in. 

Optimum stellite liner thicknesses, considering 
thermal stresses only, are given as Vi6-in. for the 100- 
round burst and Vg-in. for the 250-round burst. If barrel 
strength also is considered, the former figure remains 


unchanged, but the latter increases to 3 /io in. The 
order of the thermal stresses (compressions) at the in¬ 
terface of the Vi6-in. liner is shown in Table 1. 

These thermal stresses are accordingly significant 
in design problems, as discussed in Section 26.5.2. It 
should be remembered, however, that they exist for 
only a small fraction of the time of a round. 


7 2 3 Circumferential Strain at 

Outer Surface 

A method of determining heat input by strain meas¬ 
urement is mentioned in Section 5.3.7. The proof of 
the underlying theorem, which applies to strains in a 
heated barrel, is outlined in the present section. 

The circumferential strain at the outer surface of a 
heated cylinder depends only on the quantity of heat 
in the cylinder and is independent of the cylindrically 
symmetrical distribution of the heat, provided that 
axial shear strain is neglected, and the specific heat 
and the expansion coefficient are assumed to be inde¬ 
pendent of temperature. It is assumed that from a 
tube of inner radius a. and outer radius b, an inner 
cylindrical barrel section of radius r is removed, 
heated, and then reinserted in the outer cool section 
by a shrink fit. By equating the two formulas for the 
resulting interface radius, equation (10) is obtained. 


, Q p (r 2 + a 2 \ 

+ ra9 ~ ~E r \^r^ ~ ») 

/b 2 + r 2 , \ 

W=* + U 


= r + f-r, 


( 10 ) 


Then, if the temperature rise 0 results from a heat in¬ 
put B per unit area, equation (11) for p can be ob¬ 
tained. 


/ EqcB\ ( a(b 2 — r 2 ) \ 

p = 1 Jc) 

lending external expans 

12 ), 


(ii) 


The corresponding external expansion W e is given 
by equation (12), 


(12) 


CONFIDENTIAL 




















156 


BAND PRESSURE AND RELATED STRESSES 


which by combination with equation (11) is simplified 
to equation (13), 

w K /1D \ 

W '~~iT ]P=T (13) 

in which K is the wall ratio. The expansion W e , is 
therefore independent of r, and is equal to the expan¬ 
sion of the tube when the heat is uniformly distrib¬ 
uted. If the heat has any cylindrically symmetrical 
distribution, the proof follows from the principle of 
superposition, since the equations are linear. 


7 3 STATIC MEASUREMENTS WITH 37-MM 
AND 75-MM PROJECTILES 

731 Purpose of the Investigation 

The theoretical studies referred to in Sections 7.1.3 
and 7.1.4 were accompanied by static measurements 
with 37-mm and 75-mm projectiles. The aim of this 
investigation was 64 to arrive at an understanding of 
the physical processes occurring during and after en¬ 
graving, so that the behavior of shells during engrav¬ 
ing might be predicted for cases in which measure¬ 
ment is impossible. 

7 3 2 Apparatus and Method 

The early work consisted in setting up apparatus 
and developing methods for measuring the total 
radial load due to band pressure, and also the width 
of the pressure band, that is, the width of that por- 



Figure 3. A radial-deflection scanner. (From figure 
1 NDRC Report A-312.) 


tion of the actual rotating band on which most of the 
radial load is impressed. For a “thin” tube, with wall 
ratio 1.5, a radial deflection scanner, pictured in 
Figure 3, was employed. This method was a refine¬ 
ment of the inspection method of checking tube diam¬ 
eters by means of a 60-degree steel V-block under a 
dial gauge. The three contact points were the dial 
plunger and two spherical anvils set 120 degrees 
apart. The dial reading then gave the sum of the 
radial changes on profiles 120 degrees distant. Accu¬ 
racy of readings was improved by maintaining a con¬ 
trolled temperature. 

A section of steel tube cut from a 37-mm barrel was 
placed in the scanner and an “initial scan” with no 
pressure applied was obtained by taking dial readings 
at 0.05-in. intervals. The tube section was then re¬ 
moved from the scanner and pressure applied to the 
middle of the tube section, either by a band of known 
width, or by the rotating band of a projectile pushed 
in by a testing machine. A second scan was then 
made. Finally the tube was removed, the pressure re¬ 
leased, and final scans made to determine whether a 
permanent set had been produced in the tube. This 
was likely to have occurred with tubes having low 
wall ratios. 

Curves were then drawn plotting three times the 
radial deformation against distance along the cyl¬ 
inder. A typical graph is shown in Figure 4. Similar 
curves for known band widths were used in calibrat¬ 
ing deflections observed during the actual push tests, 
while the pressures were found by the method of Sec¬ 
tion 7.1.3. 

For a “thick” 37-mm tube, with wall ratio 2.18, 
and for sections of larger guns, a different type of 
scanner was necessary. The cylinder was supported 
horizontally in a machinist’s lathe, and a ring carry¬ 
ing a dial gauge and anvils to measure the deforma- 



Figure 4. Threefold radial deformation. (Figure 3. 
NDRC Report A-312.) 


CONFIDENTIAL 


















STATIC MEASUREMENTS 


157 


tions was so mounted that it could be moved along 
the cylinder in measured steps. 

7 3 3 Evaluation of Band Pressure 

Effective Area of Interference (EAI ). The early tests 
were carried out with solid 37-mm shot, and band 
pressure was measured at a point in the tube about 
3 in. beyond that at which engraving takes place. It 
was indicated that the pressure is small on the fore¬ 
most part of the band and that it increases toward 
the back. In analyzing the results of these tests, the 
concept of “rectangular interference ’’ was used. Since 
this quantity ignores the depth of rifling, it was later 
decided to use “Effective Area of Interference,” here¬ 
inafter abbreviated EAI. 114 This function may be de¬ 
fined as the quotient obtained by dividing the volume 
of band metal, expressed in cubic inches, that would 
be sCooped out in passing through the bore if tube 
and shot were rigid, by the bore circumference. It will 
be noted that the method of evaluating EAI differs 
according to whether band diameter is smaller or 
larger than the groove diameter. 

Band Pressure per Inch Circumference. In these 
static studies it was found convenient to use band 
pressure per inch circumference (designated by P), 
rather than pressure (per square inch). It is found by 
integrating the deformation curve over the length of 
the gun and applying the theorem on radial deforma¬ 
tion given in Section 7.1.3. The key equations are 

f wdx = g(p fl i ) /Pdx (14) 

and P = Jpdx = ^^-(p 2 — 1) JSwdx (15) 

in which 2 r 2 is the bandwidth, p the wall ratio, and p 
the pressure in pounds per square inch. 

Friction Coefficient. A third quantity of particular 
interest in these static tests is a friction coefficient, 
defined as the ratio of the axial load to the product of 
P and the bore circumference. The axial load neces¬ 
sary to move the projectile can be measured directly 
on the testing machine. In the 37-mm solid shot, fric¬ 
tion coefficients were found to vary from 0.117 to 
0.270, clustering around 0.230. 

7 3 4 Prediction of Band Pressures 

The theory of thin tubes was considered in relation 
to the observed deformations. It was found that rea¬ 


sonably good agreement could be obtained by a suit¬ 
able adjustment of parameters. 

Probably the most significant results of the study 
lie in the comparison of radial load with EAI. A linear 
relationship appears to exist for 37-mm shot, and 
reasonably good predictions of load can be made from 
EAI. It is necessary, however, to consider wall ratio 
and tightness of band seat. 114 The great variation that 
was observed in the tightness of seating of the rotat¬ 
ing bands of 37-mm shot caused a corresponding vari¬ 
ation in the apparent EAI. In a final series of tests 115 
with 75-mm AP shot and HE shell, the projectiles 
were subjected to push tests in a used 75-mm, M1897 
barrel. For the solid shot it was found that radial load 
P is proportional to EAI, as is shown in Figure 5. The 
average ratio was 78 pounds per inch circumference 
for 10 -5 sq in. EAI, but the ratio increased for very 
small bands. The main part of the pressure ap¬ 
peared to be concentrated in a narrow part of the 
rotating band. The behavior of the bands as far as 
P/ EAI was concerned was practically duplicated 
with a punch moving radially inward against the 
band. 

For the 75-mm HE shell, P/EAI was approximately 
37 for intermediate band sizes, with an increase of the 
ratio for very small bands. For very large ones, P be¬ 
came practically constant. In this connection consid¬ 
eration was given to the elastic deformation of the 
shell, and also to the firmer seating of the band dur¬ 
ing engraving. When allowance was made for these 
factors, the ratio P/ EAI was increased, even beyond 
the value for AP shot. For values of EAI in excess of 
0.011 sq in., P remained essentially constant, which 
indicated that plastic deformation of the shell wall 
had occurred. 

An extension of the study of band pressures 
in 37-mm projectiles led to interest in band design 
giving low stresses, 114 as summarized in Section 
27.4. 

Calculations of P/ EAI were made for guns of a 
number of different sizes based on firing tests at Aber¬ 
deen Proving Ground and push tests at Watertown 
Arsenal. 117 They showed that P/ EAI is independent 
of the size and design of the rotating band, except 
that this ratio is larger for very small bands. It .was 
found that in general the value of P/EAI in firing 
tests is 10 to 30 per cent higher than in push tests. 
Also, it was found that P/EAI for HE shells is 
much lower than for solid shot in small calibers, but 
that in larger calibers the values approach each 
other. 


CONFIDENTIAL 




158 


BAND PRESSURE AND REL ATED STRESSES 


90 


1C 

8 

2 80 


8 


70 


50 


a: 

s 


x 

H 


uj 60 

a 

< 

i 


UJ 

u. 

o 

X 

O 

z 

a: 

U) 

a. 

8 

S 

£t 

lil 

U. 

2 

3 

O 

tt 

O 

X 

o 

z 

OS 

UJ 

Q. 

O 

<* 

o 


40 


30 


20 


< 

5 10 
£ 





T. 

345 

341 ** 






A 






1 

• 312 

• 330 

332 • 328 0329 | 

531 * ^•336 

316 %325_1_ 





• 

• 

• 304 

• 31 ° *337 

• 323 

• 342 

309 

315 



414 



©324 
• 303 
321*305 

307^ •Iff 

109 • 

314*. M1 


B 

407 

O 

408 O 

404 

O 

409° 

04100412 

4„CP 0 04,7 

413 

4 3°a72 4250 

4 42. 

421 423 ° 

• 333 
334* • 
327 * 

• 335 

** •317 

• 308 

• 320 

318 


0401 0405 

0402 0403 

o 

406 



•322 

°428 

°427 


* 

• 75 MM AP, 

o 75MM HE, 

CURVE A 

.CURVE 8 

O 

429 















0 200 400 600 800 1000 1200 1400 

EFFECTIVE AREA OF INTERFERENCE (EA1) IN 10‘ 9 SQ IN. 

Figure 5. Band pressure as function of EAI for 75-mm AP and HE projectiles. (Figure 62, NDRC Report A-443.) 


7 3 5 The Process of Engraving 

The static test measurements were also used to 
study the process of engraving. 114 At the beginning of 
the test, that is, at zero penetration, the conic taper 
of the band rested against the forcing cone. The pro¬ 
jectile was then pushed into the tube by 0.2-in. steps, 
deformations being measured after each step. In a 


new 37-mm tube with standard forcing cone, radial 
load was shown to rise smoothly with shot travel. The 
average radial pressure P was fairly constant during 
engraving. In a moderately-eroded 37-mm tube, en¬ 
graving continued over a considerable distance. It 
was found that the radial load at a given position of 
the shot had the same value it would have in a new 
tube for a correspondingly smaller band. 


CONFIDENTIAL 

































DYNAMIC MEASUREMENTS 


159 


The axial load behaved very irregularly, depending 
on the tube. A maximum of from 110 to 200 per cent 
of the value found after engraving usually occurred 
near the end of engraving. 

Tests were made on a forcing “cone” of 90-degree 
angle. The resulting radial loads were found to be 
only one fifth to one sixth of those for a standard 
forcing cone. The rifling in this case acted like an 
axial punching tool. For purposes of comparison, the 
load for a circular punch penetrating a thick plate 
was measured as a function of penetration. The be¬ 
havior in the two cases was quite similar. 


Mechanism of Engraving 

It was concluded 114 that axial shearing is not the 
principal mechanism involved in the engraving of a 
rotating band by a forcing cone of small angle, such 
as in a Service gun. It seems that the principal effect 
is a radial pressure toward the axis of the projectile, 
resulting in plastic flow of the band material normal 
to the radius. Because of the angle of the taper this 
flow results in considerable displacement of metal 
rearward, as is shown in Figure 6. Any shearing that 



Figure 6. The three stages in the engraving of the 
rotating band of a projectile. 

occurs is only at the sides of the lands in a radial 
direction. In addition to the motion radially inward 
and axially backward, band metal also flows sideways 
from the space to be occupied by the lands to the 
space under the grooves of the rifling. 


It appears probable that the angle of the forcing 
cone does not matter as long as it is sufficiently small. 
If it approaches 90 degrees, however, then the square 
ends of the rifling can act like a punching tool in an 
axial direction, which causes shearing across the ro¬ 
tating band. 


74 DYNAMIC MEASUREMENTS DURING 
FIRING OF 3-IN. AND 37-MM GUNS 


Firings with 3-in. Gun 


Strain measurements on the external surface of the 
gun barrel were taken during the firings of the 3-in. 
gun at Carderock, as described in Section 4.4.10. 
Components of tangential strain due to gas pressure 
and band pressure may be computed from tables 
issued by Watertown Arsenal. 251 By suitable interpo¬ 
lation and extrapolation, these tables may be used for 
any wall ratio. Formulas-employed for this transfor¬ 
mation are given as equations (16) and (17), where w 
is the wall ratio (ratio of outer diameter), e t is the 
tangential strain on the outer surface of the gun, Z is 
the distance along the gun, in calibers, and the sub¬ 
scripts 1 and 2 refer respectively to the “unknown 
gun” and to the “known gun” for which values are 
tabulated. 



A tangential strain gauge was placed on the surface 
of the 3-in. gun at a point 32 in. from the muzzle. 
Curves for this gauge compiled from the Watertown 
tables are shown in Figure 7. 

In computing the band pressure, it is assumed that 
friction and acceleration produce negligible compo¬ 
nents of tangential strain, and that if the component 
due to gas pressure is subtracted from the measured 
strain, the value of the band strain is obtained. From 
this the band pressure is calculated. 

From observed data, strain was plotted against 
projectile displacement. (See Figure 8.) Gas pressure 
P was determined by assuming that the strain e 
measured for abscissas from 14 to 29 in. from the 
muzzle was entirely due to this pressure, which was 
then calculated by the Lame equation (18) where E 
is Young’s modulus and w is the wall ratio. 


2 P 

E(w 2 - 1) * 


( 18 ) 


CONFIDENTIAL 










160 


BAND PRESSURE AND RELATED STRESSES 


UJ 

tr 

Z) 

v> 

(A 

UJ 

tr 

CL 

U. 

O 


a: 

UJ 

0. 

X 

o 


400x10' 
350 x 10“* 
300 x10"* 
250x10-' 

200x10" 


CE 

UJ 


150X10" 


x 

o 


< 

CE 

\n 

UJ 


lOOxlO" 4 
50 xIO" 4 

0 




































/ 














i/ 



•COMPO 
UNIT C 

'NENT C 
;as prf 

)UE TO 
"SSURE 









/ 


— 

. COMPI 
UNIT 1 

DNENT DUE TO 
3AND PRESSUF 
1-INCH BAND ' 

3‘S MODULUS = 

IE 






/ 

/ 

7* 

s 

A 



OVER 

Y0UN< 

WIDTH 

30X10 4 






7 

/ 

/ 

/ 

/ 

\ 

\ 

\ 

\ 



POISSON'S R 

WALL RATIO 

ATIO = 

S 

0.285 

1.717 





✓ 

✓ 

'J. 



s. 

V 

N. 

} 





















PROJECTILE POSITION IN INCHES BETWEEN 8REECH END OF BAND AND MUZZLE 


Figure 7. Computed tangential strain per unit pressure on surface of 3-in. gun at 32 in. from the muzzle. (Figure 61, 
XDRC Report A-323.) 



38 37 36 35 34 33 32 31 30 29 28 27 26 

PROJECTILE POSITION IN INCHES BETWEEN BREECH END OF BAND AND MUZZLE 

Figure 8. Graphical determination of band-pressure strain at gauge 32 T (round 43) of 3-in. gun. (Figure 62, NDRC 
Report A-323.) 


Assuming linear pressure variation, these values of 
P were extrapolated to projectile positions 29-37. 
Multiplying them by the corresponding strains per- 
unit pressure shown in Figure 7, the component of 
strain due to gas pressure resulted, as shown by the 


dotted curves in Figure 8. This was then subtracted 
from the measured tangential strain, to give strain 
due to band pressure. Band pressure in pounds per 
square inch was found by dividing the peak value of 
this curve by that of the dashed curve in Figure 7. 


CONFIDENTIAL 

















































































STATIC MEASUREMENTS 


161 



PROJECTILE POSITION IN INCHES 
BETWEEN BREECH END OF BAND AND MUZZLE 

Figure 9. Graphical determination of band-pressure 
strain at gauge 90 A (round 61) of 3-in. gun. (Figure 42, 
Third Carderock Report, A-460.) 

A somewhat similar method was used for comput¬ 
ing band pressure at gauge 121.2 T, placed only 4.6 
in. from the origin of rifling. However, the gas pres¬ 
sure was derived from that at the breech, using a dif¬ 
ferent assumption as to pressure distribution. 

The results of these measurements indicate a band 
pressure, for a 100 per cent charge round, of from 
20,000 to 30,000 psi near the muzzle, compared with 
a pressure of from 60,000 to 70,000 psi near the origin. 
For reduced-charge rounds the latter pressure shows 
little variation from that for the 100 per cent charge, 
but the pressure farther along the barrel is decreased 
to only about 40,000 psi. These band pressures exceed 
the yield point of cold-worked copper, the material of 
the rotating bands used in the firings. 

An interesting special measurement was that of 
strains in the vicinity of a hole in the gun bore which 
had been drilled at right angles to the axis for the in¬ 
troduction of a pressure gauge. Eleven strain gauges 
were cemented to the surface of the gun barrel around 
the hole, one of them being an axial gauge at its edge. 
The typical curve of axial strain illustrated in Figures 
8 and 9 was modified, the tension peak being chopped 
off about half way up by a jagged line which corre¬ 
sponded to the time the rotating band was passing 
under the hole. Axial gauges 2 in. from the hole’s edge 
showed this phenomenon to a lesser degree. This re¬ 
duction in strain is attributed to local reduction in 
band pressure over that portion of the band, which 
passes under the hole. 


7 4 2 Studies With Army 37-mm Field Gun 

High-speed strain-recording equipment for use in 
general stress research has been developed in recent 
years at the Massachusetts Institute of Technology. 
It was tested at Watertown Arsenal 249 in firings of a 
37-mm field gun, and characteristic firing strain rec¬ 
ords were obtained. 166 

Eight strain gauges of the same type as was used 
later on the 3-in. gun at Carderock were used in com¬ 
bination with a special three-element film-recording 
oscillograph. Two gauges were placed tangentially 
over the powder chamber, and pairs of tangential and 
axial gauges were placed near the origin of rifling, 
midway down the barrel, and near the muzzle. The 
firing-strain records of the two gauges near the pow¬ 
der chamber rose smoothly to maxima and then de¬ 
creased gradually. Gauges 3, 4, 5, and 6 exhibited 
curves similar to those of Figures 1 and 2. 

Gauge 7 indicated some vibrational effect, but this 
was much more pronounced for gauge 8, the tangen¬ 
tial gauge near the muzzle. These vibratory strains 
developed earlier and were greater in amplitude when 
the projectile had a muzzle velocity of 2,650 fps than 
when the velocity was 2,520 fps. At 2,000 fps the 
effect was hardly noticeable. 

Later a detailed study of band pressures in these 
firings was made. 327 The computed band pressure 
reached a maximum of 78,100 psi. This value is of the 
same order of magnitude as that found for the 3-in. 
gun used in experiments described in Section 7.4.1, 
when the bore was greased. 

Several types of theoretical analysis were intro¬ 
duced in this study. The first is a modified shear anal¬ 
ysis, which gave good results for band pressures dur¬ 
ing firing when taken in conjunction with oscillograph 
records of firing strains. When this distortion was 
later combined with bending, agreement with experi¬ 
ment was again good, except for maximum stresses, 
which checked better with full shear analysis. 

Two general methods were employed in interpret¬ 
ing the firing data. The first method 339 assumes that 
the rotating band acts like a narrow band of high 
pressure moving down the barrel, followed by the 
powder pressure. The crest of the transient would 
occur at the instant the center of the rotating band 
passes the gauge location, if powder pressure is neg¬ 
lected. For gauges 6, 12, 15, 18, and 24 in. from the 
muzzle of the 3-in. gun this assumption gave excel¬ 
lent agreement. 39 Gauges farther from the muzzle 
gave readings not readily interpreted because of other 


CONFIDENTIAL 



















162 


BAND PRESSURE AND RELATED STRESSES 


factors involved. The second method 327 is a modifica¬ 
tion 341 - 346 of the relaxation procedure of analysis in¬ 
troduced by the British, which is described in Section 
27.2.3. 

75 DYNAMIC MEASUREMENTS DURING 
FIRING OF CAUIBER .50 MG BARRELS 

751 Experimental Conditions 

Strain measurements were made during the firing 
of four caliber .50 aircraft machine gun barrels. Two 
of the barrels were fitted with 9-in. liners of Stellite 
No. 21, as described in Section 22.3. The third was a 
nitrided, chromium-plated barrel, in which the plat¬ 
ing was slightly tapered toward the muzzle, as de¬ 
scribed in Chapter 23. The fourth barrel was a regular 
steel one. 

Four types of ammunition were used: API-M8, a 
common combat round; AP-M2, the ammunition or¬ 
dinarily used in the erosion tests of caliber .50 ma¬ 
chine gun barrels; Tr-Ml, which previous experiences 
had shown to give the least erosion; and Tr-MlO, 
which previously had given the most erosion in caliber 
.50 aircraft barrels. (See Chapter 23.) The largest and 
smallest diameter AP-M2 bullets were especially fired 
for purposes of comparison. 

7.5.2 Measurements and Computations 

Strains were measured with resistance gauges ce¬ 
mented circumferentially at chosen locations on the 
outer surface of the barrel, the results being recorded 
by the oscillographs described in Section 4.4.2. Max¬ 
imum tangential stresses were then tabulated. Start¬ 
ing at the origin of rifling, these ranged from 10,000 
psi up to values approaching 20,000 psi at 10 in. from 
the breech, then decreased as the bullet approached 
the muzzle. This was not characteristic of stresses 
when oil was present in the bore, however. 

In addition to the determination of tangential 
stresses, calculations were made for five other quan¬ 
tities : 


1. Maximum bore stresses, computed from the 
outer surface stresses by the Lame formula given in 
equation (19 

Pb = P 0 wl ~ H ( 19 ) 

in which P B and P 0 are the bore and surface stresses, 
and w is the wall ratio. 

2. Gas pressures, similarly computed from the 
strain-gauge records, after the band pressure peak 
had passed. 

3. Band pressure, computed by a method de¬ 
veloped at Watertown Arsenal. 249 The band width 
was assumed to be 1.6 calibers. 

4. Bore expansion, e ta m, calculated from the wall 
ratio w and the maximum external tangential strain, 
e tm , according to the Lame theory as expressed in 
equation (20) in which /x, the Poisson ratio, is taken 
to be 0.3. 249 

e ta J l -.^ + (1 + ^ (20) 

5. Heat input, by the method discussed in Section 
5.3.7. (See also Section 7.2.3.) 

7,5,3 Discussion of Results 

In a nitrided and chromium-plated barrel with ta¬ 
pered plating, the band-pressure peak disappeared at a 
point between 14 and 24 in. from the breech; in the 
stellite-lined barrel this occurred around the end of 
the liner, at about 12 in. from the breech. 

Band pressure, when present, was greater than the 
gas pressure. When the bore was oily, these band 
pressures, except at the origin of rifling, were higher 
and lasted for a greater distance down the barrel. 

Outer surface stresses on the chamber section of a 
lined barrel were smaller than on a monobloc barrel. 
The more erosive ammunitions produced the greatest 
strains. Maximum gas pressures ranged from 45,000 
to 53,000 psi and maximum band pressures from 
51,000 to 75,000 psi. The bore expansion was of the 
order of from 0.001 to 0.002 in. 


CONFIDENTIAL 




Chapter 8 

EXTERIOR BALLISTICS OF HYPERVELOCITY PROJECTILES 

By A. H. Stone a 


81 INTRODUCTION 

H ypervelocity projectiles have many advan¬ 
tages, as far as exterior ballistics is concerned, as 
indicated in Chapter 1, provided that they can be de¬ 
signed so as to satisfy two conflicting requirements. 
These are: (1) stability (so that the yaw is kept small), 
and (2) high ballistic coefficient (i.e., low air resist¬ 
ance). Calculations, confirmed by firings, show that 
these requirements can be met satisfactorily by sub¬ 
caliber projectiles and evaluate their advantages. 

Much research was involved in these developments. 
Simplified methods have been devised for predicting 
the behavior of projectiles in the hypervelocity range. 
The stability and the optimum proportions of sub¬ 
caliber projectiles have been investigated for a selected 
minimum value of the stability factor. Since experi¬ 
mental ballistic data at high velocities were lacking 
for a series of projectiles having a systematic varia¬ 
tion of dimensions, a theoretical attack was made on 
the behavior of the drag and the stability at high 
velocities through the application of the dynamics of 
compressible fluids. The theory also makes possible 
an estimation of the effects of varying ogive lengths. 


in particular on subcaliber projectiles, the results ob¬ 
tained are of value for exterior ballistics in general. 

82 FLIGHT CHARACTERISTICS OF 
HYPERVELOCITY PROJECTILES 

8,2,1 Hypervelocity and Energy Loss 

Hypervelocity projectiles differ in their exterior 
ballistics from standard projectiles in two respects: 
(1) there is an effect caused purely by their higher 
velocity; (2) the projectiles themselves are generally 
different from the standard type, and in the cases of 
greatest interest are subcaliber. 

The first effect can be described by saying that a 
given projectile overcomes air resistance more effi¬ 
ciently at hypervelocity than at conventional speeds. 
The drag force d resisting the motion of a projectile is 
expressible as 176>509 

d = pK D D 2 U 2 , (1) 

where p is the air density in pounds per cubic foot, D 
is the diameter of the projectile in feet, and U is its 
velocity in feet per second. The dimensionless coeffi- 


Table 1. Remaining velocities (RV) and energy losses (EL) of Gg-type projectile traveling over various ranges. 


Range (yd) 

Muzzle 500 1,000 1,500 2,000 

velocity RV EL RV EL RV EL RV EL 

(fps) (fps) (per cent) (fps) (per cent) (fps) (per cent) (fps) (per cent) 

5,000 4,710 12 4,430 22 4,160 31 3,890 40 

4,000 3,730 13 3,470 25 3,210 35 2,960 45 

3,000 2,750 16 2,510 30 2,270 43 2,040 54 

2,000 1,780 21 1,570 38 1,380 52 1,200 64 


The analysis of the results of firings is greatly simpli¬ 
fied b}' the development of a photographic method of 
trajectory determination. Finally, the effects of vari¬ 
ous design features have been studied by systematic 
firings carried out with the cooperation of Aberdeen 
Proving Ground. 

While attention was focused on hypervelocities and 


a Geophysical Laboratory, Carnegie Institution of Washing¬ 
ton. (Present address: Cambridge University, Cambridge, 
England.) 


cient K d is, for any one projectile, a function of the 
velocity (or, more accurately, of the ratio U/ai where 
ai is the velocity of sound). In the supersonic range, 
K d is found to decrease gradually from a maximum in 
the neighborhood of the velocity of sound. Figure 1 
illustrates a typical case. [Systematic data are lack¬ 
ing at velocities above 4,000 fps and extrapolation is 
used; but theoretical considerations (Section 8.5) show 
that this is not likely to cause much error.] Thus 
while increasing U increases the drag it also decreases 
the ratio of the drag to the kinetic energy of the pro- 


CONFIDENTIAL 


163 








164 


EXTERIOR BALLISTICS OF HYPERVELOCITY PROJECTILES 



Figure 1 . The drag coefficient Kd for a projectile of 
type G8. (This figure is based on Figure 1 of Ballistic Re¬ 
search Laboratory Report No. 409, Aberdeen Proving 
Ground.) 

jectile. Consequently, in traveling over a given range, 
a projectile at hypervelocity retains not only a greater 
kinetic energy, but a greater proportion of its original 
kinetic energy as compared with the same projectile 
at a lower but still supersonic velocity. This effect is 
seen in Table 1, which shows the remaining velocity 
(RV) and the percentage of the kinetic energy that 
has been lost (EL) in a typical case, for various muz¬ 
zle velocities and ranges. The table has been com¬ 
puted for a projectile b of type Gs and ballistic coeffi¬ 
cient 1.0 when fired horizontally. 

8 2 2 Range and Time of 

Flight at High Speeds 0 

Hypervelocity has a further advantage to the ex¬ 
terior ballistieian, in that it permits simplification of 
computations, particularly over ranges short enough 
for the curvature of the trajectory to be neglected (as 
is usually the case with hypervelocity projectiles in 
practice). If correction factors for variations in air 
density and air velocity are omitted, the retardation 
(— dU/dt) produced by the drag of the projectile can 
be expressed 509 by equation (2), 

_dU_ pUF(U) 
dt ~ C ’ 


b The Gs type of projectile is characterized by a cylindrical 
base and a long ogive (radius of curvature: 5 to 10 calibers). 
It is the standard British projectile form. 

0 Based on a memorandum, not available for general circula¬ 
tion, from C. W. Beck and K. F. Herzfeld, Catholic University, 
to the Chief of Division 1, NDRC, 1942. See also pp 67-75 of 
NDRC Report A-234. 42 


in which for a given standard projectile shape, F(U) 
is a definite function of the velocity (or, more exactly, 
of the Mach number, U/ai), and C, the ballistic co¬ 
efficient, is defined by 

C = m/iD\ (3) 

where m is the mass of the projectile and i is the 
“form factor” 172488 —a number usually close to 1 that 
corrects for the deviation of the actual projectile 
shape from the standard one to which F strictly ap¬ 
plies. In equation (3) and subsequent equations based 
on it, the projectile diameter D is measured in inches, 
instead of in feet, as in equation (1). For ranges that 
are short enough for the curvature of the trajectory 
to be neglected, equation (2) leads to 

B= _c/' P _«L (4a) 

L Ju oP F(uy 

t - cf U <EL (4b) 

c Ju 0 pUF(uy ( ’ 

where the projectile velocity falls from Uo to U in 
traversing a distance R in a time t. Now, over the 
hypervelocity range, the function F(U) is very 
“smooth,” partly because it is obtained by extrapola¬ 
tion, and can be approximated sufficiently well by the 
linear expression 

F(U) =AU + B, (5) 

where A and B are suitably chosen constants, easily 
determined graphically. For example, it has been cal¬ 
culated that in the range 3,800 ^ U ^ 6,000, if 
A = 0.0001 and B = 0.0123, the error is less than 
one-third percent. Here the projectile* 1 is of type Gi, 
and the velocity is measured in feet per second. In a 
wider velocity range, 2,400 ^ U ^ 6,000, if A = 
0.00009 and B = 0.06, the error is less than 3 per 
cent. 42 The substitution of AU -f B for F(U) in equa¬ 
tions (4) enables the integration to be carried out ex¬ 
plicitly. Thus 

R (in yards) = In (6) 

t (in seconds) = J In ( 7 ) 

and the relation between t and R can be obtained. 42 
Equations (6) and (7) also enable C to be determined 

d The Gi type of projectile is characterized by a square base 
and a blunt ogive (radius of curvature: 2 calibers). It was one 
of the types fired at Gavre. 


CONFIDENTIAL 


















STABILITY OF SUBCALIBER PROJECTILES 


165 


from a firing of the projectile. For this purpose it is 
found that a single firing over a long trajectory gives 
more accurate results than several firings over shorter 
distances. 

8 2 3 Subcaliber Projectiles 

The effect of replacing a standard projectile by one 
of subcaliber can now be considered. It is evident 
from equation (3) that scaling down a full-caliber pro¬ 
jectile to one of subcaliber tends to decrease the bal¬ 
listic coefficient C, because the mass varies as the 
cube of the caliber. Thus, a comparison between the 
two types of projectile shows that 

Ci/G ^ Di/D 2 , 

where the subscript 1 refers to the subcaliber projec¬ 
tile, and the subscript 2 refers to the one of full cal¬ 
iber; hence, the retardation of the subcaliber projec¬ 
tile will be at least D%/ A times the retardation of the 
full-caliber projectile. This effect is accentuated by 
the fact that, in practice, requirements of stability 
(see Section 8.3) may oblige the subcaliber projectile to 
be shorter in proportion, which results in a further 
lowering of its mass and in an increase in the form 
factor. 

To investigate how far this effect may offset the 
advantage of the higher velocity of the subcaliber 
projectile, calculations have been carried out for the 
least favorable case, namely, that of the sabot-pro¬ 
jectile (described in Chapter 29), for which the kinetic 
energy of the sabot is wasted. Using the method just 
described, it has been shown 42 that over short ranges 
the ratio of the times of flight t\ and 4 for a typical 
sabot-projectile and a full-caliber projectile, respec¬ 
tively, is approximately given by equation (8), 

h _ Ci In {[F(Ui) - Be AR / Ci ]/AUi] {0 , 

k C 2 ln {[F(U 2 ) - Be AR /c>]/AU 2 \’ 

in which R is the range, L\ and U* are the muzzle 
velocities of the two projectiles, and A and B are as in 
equation (5). From equation (8) the times of flight of 
a 57/75-mm sabot-projectile and a 75-mm full- 
caliber projectile fired from the same gun have been 
calculated for various ranges as shown in Figure 2, 
using Ci = 2.0, C 2 = 2.6, Ui = 2,805 fps, and U 2 = 
2,050 fps. 

It will be seen that up to a range of 5,000 ft, the 
time of flight of the sabot-projectile is only about 
three-quarters of the time of flight of the correspond- 



R IN FEET 

Figure 2. Ratio of time of flight of 57/75-mm sabot- 
projectile to that of a full-caliber projectile fired from the 
same gun to various ranges. (This figure has appeared 
as Figure 18 of NDRC Report A-234.) 

ing full-caliber projectile. Furthermore, over this 
range the two projectiles have nearly equal kinetic 
energies; and since the sabot-projectile has its energy 
concentrated over a smaller area, it has the advan¬ 
tage in armor-piercing ability, as discussed in Section 
9.2.2. 

The above calculations refer to an extreme case. In 
practice a compromise is often made by increasing the 
mass of the subcaliber projectile, usually by giving 
it a core of tungsten carbide. This results in a lower 
initial velocity but a more favorable ballistic coeffi¬ 
cient, and, moreover, tungsten carbide leads to im¬ 
proved armor penetration, for the reasons given in 
Section 9.2.2. Another reason for this compromise is 
the need for stability, which will now be considered. 

83 STABILITY OF SUBCALIBER 
PROJECTILES 

8,3,1 General Theory 

The requirement of stability is crucial for subcaliber 
projectiles, especially sabot-projectiles. Accordingly 
the following account begins by briefly reviewing the 
general considerations involved. 

In general, a spinning projectile in flight does not 
point exactly in the direction of its motion, but yaws 
in an oscillatory way; and it is important that the 
yaw should be damped out, or at least should not in¬ 
crease. The projectile is then said to be stable. 

In the first instance, the situation may be simpli¬ 
fied by disregarding the more complicated features of 
air resistance and considering only a short portion of 
the trajectory, so that the retardation of the projec¬ 
tile and the effect of gravity can be disregarded. The 
air resistance acting on the projectile is then equiv¬ 
alent to a single force /, which can be resolved into 
two components: (1) the axial drag d acting along the 
axis of the projectile, and (2) a force L perpendicular 
to the axis, acting through a point P on the axis 


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166 


EXTERIOR BALLISTICS OF HYPERVELOCITY PROJECTILES 



Figure 3. Simplified representation of the forces that 
act on a spinning projectile in flight. The air resistance is 
considered as a single force having two components: the 
axial drag d acting along the axis of the projectile and a 
force L perpendicular to the axis, acting through the 
center of pressure P. For a spin stabilized projectile, P is 
in front of the center of gravity G. The angle of yaw is e. 


called the center of pressure (see Figure 3). For spin- 
stabilized projectiles, the center of pressure P is in 
front of the center of gravity G. 

The moment tending to overturn the projectile is 
then M = LD (h — g ), where g and h are the dis¬ 
tances (in calibers) from the base A to the centers of 
gravity and pressure, respectively, and D is the cal¬ 
iber of the gun. 

It is found that L/e and h are nearly independent 
of the yaw e (measured in radians), provided e is not 
too large; hence 

M = LD(h-g) = M e, ' (9) 


Avhere \i can, for the moment, be regarded as constant. 
The condition for stability 184 199,509 is then s > 1, 
where s, the stability factor, is defined by 

AW 2 nrvi 

s = w (10) 


in which A is the axial moment of inertia of the pro¬ 
jectile, B is the moment of inertia about the trans¬ 
verse axis through the center of gravity, and N is the 
axial spin of the projectile (in radians per second). If 
the condition that the stability factor is a constant 
greater than unity is satisfied, the motion of the nose 
of the projectile as seen from the center of gravity 
will be epicyclic. For the simplified set of forces shown 
in Figure 3, the motion is that of a simple epicycle, 
which is the path traced by a point attached to a fixed 
point by two jointed links, each of which rotates 
uniformly, as illustrated in Figure 4A. 


8 3 2 Stability Under Actual Conditions 

If now the full effects of the air resistance are al¬ 
lowed for, the situation becomes more complicated. 
There are forces and couples present caused by the 


Magnus effect, by the frictional resistance to the spin 
of the projectile, and even by the yawing oscillations 
themselves. The complete theory 197,444 shows that if 
s > 1 the nose of the projectile oscillates in much the 
same way as before, but the lengths of the epicyclic 
arms are exponentially damped, that is, are multi¬ 
plied by factors of the form e~ CR , where R is the range 
and c is a parameter that is nearly constant. This 
damping causes the path described by the nose of the 
projectile as seen from the center of gravity to be that 
shown in Figure 4B. 509 




Figure 4. Epicyclic motion of the nose of a projectile: 
A. Retardation and effect of gravity disregarded. B. Full 
effects of air resistance considered. 


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STABILITY OF SUBCALIBER PROJECTILES 


167 


Moreover, it is possible for this damping to be neg¬ 
ative, so that the yaw can actually increase, even if 
s > l. e The condition for positive damping of the yaw 
is found 1 to be 


jrK„ - ~ (K a - 2 Kr) 

where V = -, - -——r- (11) 

-^ Kl + - {Ka _ 2Kt), 

in which D is the diameter of the projectile, m its 
mass, A and B the moments, and K h ,Kl, K a , K t , are 
various dimensionless coefficients, all nearly constant. 
Equation (11) is more restrictive than the condition 
s > 1. The coefficients involved are difficult to meas¬ 
ure ; but, using the best estimates® available, s should 
be at least 1.2 for projectiles of conventional type if 
the yaw is to be damped at all, and should preferably 
be at least 1.5 for rapid damping, namely, a shrink¬ 
age of both epicyclic arms to one-half of their initial 
value in 3,000 calibers of travel. A value of s of 1.3 
may be taken as a minimum. For unusually long pro¬ 
jectiles, e.g., rockets, the limit for s is larger; and 
very long projectiles may possibly be incapable of 
stabilization by spin. 

8 3 3 Stability Over Long Arcs of the 
Trajectory 

The foregoing applies only to a short arc of the 
trajectory. Over longer arcs, the variation of \x must 
be taken into account. The “moment coefficient” m 
can be expressed 509 b}^ equation (12), 

m = pK m D 3 U 2 = pK N D\h-g)U 2 (12) 

(where K N = L[pD 2 U 2 e ); here, as before, p denotes 
the air density, U the velocity, and D the caliber of 
the projectile. The dimensionless coefficients Am and 


e Personal communication from I. E. Segal, Ballistic Re¬ 
search Laboratory, Aberdeen Proving Ground, as given in his 
memorandum to R. N. Thomas, September 28, 1944. An 
example was given in that memorandum. 

f This equation is based on one published by Kent and 
McShane 199 (equation 31). A substantially equivalent condi¬ 
tion, in the notation of Fowler 444 is: 

c> (* + *)« 

4(fc + 7 ) (ft - 7 ) 



1.0 1.5 2.0 

U/cq 

Figure 5. The coefficient Km in equation (12) as a 
function of the velocity for a projectile of type J. (Re¬ 
printed by permission from Elements of Ordnance by T. 

J. Hayes, published by John Wiley and Sons, Inc. This 
figure is based on Figure 223 in that work.) 

Kn are, for a given projectile shape, functions of U (or 
rather U/ch);K M does not vary greatly in the relevant 
range (see Figure 5). 509 Along the trajectory, U de¬ 
creases faster than the spin N ; hence, from equations 
(10) and (12), s increases along the trajectory. Thus if 
a projectile is initially stable, it will remain so. 

Equation (12) shows also that the stability of a pro¬ 
jectile depends on the weather. A projectile fired into 
denser (e.g., colder) air than standard will have a 
larger moment coefficient, and hence a smaller sta¬ 
bility factor, than when fired under standard condi¬ 
tions. To allow for this, and to provide a margin of 
safety, the stability factor of a projectile when tested 
under standard conditions should be at least 1.5. 

8 3 4 Stability As Affected 

By Special Conditions 

Under special circumstances, this limit for s may be 
modified. 


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168 


EXTERIOR BALLISTICS OF HYPERVELOCITY PROJECTILES 


Small Initial Yaw 

If the initial yaw can be kept very small, a lower 
value of s may be used; even if the yaw increases at 
first, the increase of s along the trajectory may lead 
to damping of the yaw before it reaches an unduly 
large value. This consideration applies to most de¬ 
formable projectiles (Chapter 29) and to pre-engraved 
projectiles with wide rotating bands (Section 27.3) 
since both types acquire very little yaw in the gun. 

Projectile Fired From a Plane 

However, a larger value of s may be needed to 
moderate the yaw of a projectile fired from an air¬ 
plane, 174,193 where there may be a large effect caused 
by head wind and crosswind combined. 

Firing Ahead. The head-wind effect on stability is 
at its worst when the fire is straight ahead, for then 
the initial velocity relative to the air is increased from 
U to U + w (where w is the velocity of the plane) 
without a compensating increase in the spin, N. From 
equations (10) and (12) the value of e must be multi¬ 
plied by U/(U + w) 2 . In a representative case, w 
= 300 fps, U = 2,500 fps, hence the stability factor is 
decreased by nearly 20 per cent from its value when 
the projectile is fired from rest. 

Firing Abeam. The crosswind effect is at its worst 
when the fire is abeam. The projectile will then have an 
initial yaw of approximately w/U (radians) and a first 
maximum yaw of approximately w/U\/\ — 1/s. 174 
In the case just cited, with s = 1.5, the yaw is about 
12 degrees, which is undesirably large. An increase in 
s is desirable on this account, to accelerate the damp¬ 
ing of the yaw. (It also decreases the first maximum 
yaw, but only slightly.) The way in which this damp¬ 
ing depends on $ is given by equation (13), 193 

(13) 

where e is the mean yaw, 6 0 the initial yaw, R the dis¬ 
tance traveled, p the air density, a is a constant (the 
“damping constant”) and a' = D 2 Kd/2(s — 1 )m. 
This equation is based on the assumption that the 
damping rates on the two epicyclic arms of Figure 4B 
are equal. 

Thus a satisfactory value of the stability factor for 
a projectile to be fired from an airplane is about 1.7. 

Effect of Hypervelocity. Finally, one may inquire 
into the effect of hypervelocity as such on stability. 
As Figure 5 suggests, the indications are that K M 


should be slightly smaller in the hypervelocity range 
than at standard muzzle velocities. This means that a 
given projectile has a higher stability factor at hyper¬ 
velocity than at standard velocities, in other words, 
that hypervelocity has to some extent a stabilizing 
effect. However, though systematic experimental 
data are lacking in the hypervelocity range, theoret¬ 
ical considerations (which are dealt with in Section 
8.4.6) indicate that this effect is small at best, and 
may be reversed at still higher velocities. Thus on fir¬ 
ing from a stationary gun, the stability factor is sub¬ 
stantially independent of the muzzle velocity. 

^lowever, in connection with fire from airplanes, 
hypervelocity has an obvious advantage, since in effect 
it reduces the relative importance of the velocity of 
the airplane. For example, doubling U would halve 
the effects considered in the preceding subsection. 

Summary. To sum up, an acceptable value of the 
stability factor is in general 1.5, but it can be lower if 
the initial yaw is kept very small, or if the projectile 
is not to be used under conditions of high air density, 
and it should be higher if the projectile is to be fired 
from an airplane, though this last effect is less impor¬ 
tant for hypervelocity projectiles. 

8.3.5 Methods of Obtaining Stability 

From equations (10) and (12), 

A 2 N 2 

s = 4f,Ks-DHh~g)U-rr (14) 

This expression makes obvious several general 
methods 14 for modifying the design of a spin-stabi¬ 
lized projectile so as to increase its stability. 

1. The twist of the rifling in the gun can be in¬ 
creased. This increases s very effectively (by increas¬ 
ing N), but was virtually excluded during World War 
II by the great desirability of retaining standard 
guns, a consideration that might still have weight in 
the future with new guns for the sake of using stand¬ 
ard projectiles. 

2. The projectile can be shortened, thereby increas¬ 
ing A 2 /B. This has the drawback of decreasing the 
ballistic coefficient defined by equation (3). 

3. The density of the projectile can be increased. 
Here both A 2 /B and the ballistic coefficient are in¬ 
creased. There is the disadvantage that the increased 
mass leads to some loss of muzzle velocity. 

4. A compromise version of the preceding methods 
is possible whereby the projectile can be given a core 
of denser material, such as tungsten carbide. This has 


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STABILITY OF SUBCALIBER PROJECTILES 


169 


the advantage that the optimum combination for 
armor penetration can be selected. 

5. If a core is used, it may (if geometrically possible) 
be set forward. This moves the center of gravity closer 
to the center of pressure, that is, it decreases ( h-g ) 
and so increases s without affecting the ballistic co¬ 
efficient. 

6 . Finally, fins or tail surfaces may be used to move 
the center of pressure closer to the center of gravity. 
As has been pointed out, 14,42 this method would be 
especially suitable for sabot-projectiles. It has been 
tried by the Germans and deserves further investi¬ 
gation. 


8 3 6 Stability of Deformable Projectiles® 

Consider first a deformable projectile which, after 
deformation, is equivalent to a scaled-down version 
of the full-caliber projectile. We use the subscript 0 to 
refer to the deformable projectile before deformation, 
the subscript 1 to refer to it after deformation, and 
the subscript 2 to refer to the full-caliber projectile. 
From equations (10) and (12), 


Aftfh/Utf 

IpKmDJBS 


(15) 


and since the conservation of angular momentum 
shows that A\Ni — A 0 N 0 (if frictional losses are, for 
the moment, neglected), this can be written 


Similarly 


(Ap/A i) 2 Ai 2 (Nq/Ui ) 2 
IpKmDSB! 

_ A 2 2 {N 2 /U 2 ) 2 
82 4 pK M D 2 *B 2 


(16) 

(17) 


Now, No/Ui — N 2 /U 2 , both being determined by 
the pitch of rifling of the gun. Further, if both pro¬ 
jectiles have the same density, Ai/A 2 = A 5 /A 5 = 
Bi/Bi. Hence, / . \,/ n \, 

i=(£)(!) <■*> 


Since A 0 > A i, while A < D 2 , the factors on the 
right tend to cancel. For example, in the 57/40-mm 
deformable projectile described in Section 30.2, 
Ap/Ai = 1.35, Di/D 2 = 40/57, hence «i/s 2 = 0.90. 
Here the deformable projectile would be nearly as 
stable as the full-caliber projectile, except that fric¬ 
tional losses have been so far neglected. The friction 


g Compare Chapter 30. In that chapter the adjective 
“skirted” is used rather than “deformable.” 


in the muzzle adaptor will change U\ only a little, 
since it is offset by the continued pressure of the pow¬ 
der gases; but it will slow down the angular spin. Ex¬ 
periments indicate 128 that in a new adaptor about 10 
per cent of the angular momentum is lost in this way. 
(In an old adaptor the loss rises to 25 per cent.) Thus 
AiNi is equal to only (0.9) A 0 No, and equation (18) 
must be replaced by approximately 



for a new tube, the constant factor falling to 0.5 in a 
worn tube. 

The full-caliber projectile ordinarily has enough ex¬ 
cess stability for the deformable projectile to be suf¬ 
ficiently stable, particularly since deformable pro¬ 
jectiles ordinarily have small initial yaws. Further¬ 
more, attention has been concentrated on deformable 
projectiles with tungsten carbide cores; and, as was 
pointed out in Section 8.3.5, this feature increases s, 
and may be made to increase it still further by ad¬ 
vancing the core. This effect is borne out, on the 
whole, by the data 456 shown in Table 2. 


Table 2. Stability factors of some 57/40-mm deform¬ 
able projectiles. 


Type 

No. 

Distance from rear 
of core to rear of 
container (in.) 

Distance of center of 
gravity from rear of 
container (in.) 

Stability 

factor 

1932 

1.50 

2.70 

2.3 

1936 

1.44 

2.66 

4.9 

1938 

1.37 

2.61 

4.8 


Thus there is no difficulty in obtaining sufficient 
stability for deformable projectiles without undue 
sacrifice of muzzle velocity or ballistic coefficient. 

8 3 7 Stability of Sabot-Projectiles 14 - 42 

As before, we begin by comparing the stability 
factors of a standard full caliber projectile and a 
scaled-down sabot-projectile. Using subscripts 1 and 
2 to refer to the sabot-projectile and the full-caliber 
projectile, respectively, equations (15) and (17) are 
again applicable. Suppose the projectiles have densi¬ 
ties Ai and X 2 ; then A\A 2 = \\D\ /\ 2 D 2 = Bi/B 2 . In 
the present case, Ni/Ui = N 2 /U 2 , both being equal 
to 2tt/7) D 2 where r; calibers per turn is the pitch of 
the rifling. Hence 



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170 


EXTERIOR BALLISTICS OF HYPERYELOCITY PROJECTILES 




Figure 6. The ratios of muzzle velocities (U 1 /U 2 ), ar¬ 
mor penetration (AP), and efficiency (proportion of 
work done by powder which goes into projectile) rela¬ 
tive to the standard, as functions of Di/D 2 , for (A) a 
steel sabot-projectile and (B) a tungsten carbide sabot- 
projectile. (This figure is based on Figures 3(a) and 4(a) 
of NDRC Report A-88.) 


If the densities are the same, the requirement of 
stability imposes a severe limitation on the reduction 
A/A, particularly since sabot-projectiles tend to 
have appreciable initial yaws, and hence should have 
stability factors of at least 1.5. This difficulty can be 
overcome in part by choosing a standard projectile 
with a high stability factor s 2 for a model, and by in¬ 
creasing the density. 

The validity of equation (20) is confirmed by some 
firings of 75/57 mm and 75/40 mm projectiles, 42 100 in 
which, however, the subcaliber projectile was simply 
a standard projectile of lower caliber, and thus was 
not designed for optimum results. 

To explore the possibilities more fully, calculations 
were undertaken 14 to determine the best scaled-down 
projectile to be used in the 37-mm cannon for sabot- 
projectiles made of (1) steel and (2) tungsten carbide 
(for which X 1 /X 2 is nearly 2). In each case, the value of 
the scaling-down ratio A/A was chosen to give the 
greatest armor penetration at point-blank range con¬ 
sistent with a stability factor of at least 1.5. Since the 
stability factor of the standard AP shot in this gun 
is 3.0, equation (20) shows that A/A cannot be less 
than l/\/2 = 0.71 in case (1), nor less than Yl in case 
(2). The analysis leads to the results exhibited in 


Figure 6, which shows that these ratios are in fact 
almost exactly those that give maximum penetration. 

The effect of modifying the body length of the pro¬ 
jectile was also considered, with the final result that 
the sabot-projectiles giving optimum armor penetra¬ 
tion in the 37-mm cannon were found to be as follows: 

1 . Sabot-projectiles of steel should be scaled down 
from the standard 37-mm AP projectile in the ratio 
A/A = 0.7; thus the subcaliber is 26 mm. The pro¬ 
portions of the projectile are then about the same as 
those of the standard, except that the body is 10 per 
cent shorter. The muzzle velocity is 40 per cent higher 
than that of the standard, and the armor penetration 
is increased by 18 per cent. The ballistic coefficient is 
lower than that of the standard projectile; thus the 
improvement applies only to short ranges. 

2 . Sabot-projectiles of tungsten carbide should be 
scaled down from the standard 37-mm AP projectile 
in the ratio A/ A = 0.5. The proportions are the 
same as for standard projectiles; the muzzle velocity 
is increased by 33 per cent, and the armor penetra¬ 
tion by 50 per cent. Furthermore, the ballistic coeffi¬ 
cient is the same as that of the standard projectile, so 
that the reduction in time of flight, and the gain in 
velocity and armor penetration apply throughout the 
trajectory. If tungsten carbide is used merely in the 
form of a core, the results are intermediate. 


84 MOTION OF A SLIGHTLY YAWING 
CONE MOVING AT SUPERSONIC SPEEDS h 

841 Idealization 

A purely theoretical determination of the effects 
produced upon the flight of a projectile by yaw was 
carried out in a case which, while necessarily ideal¬ 
ized, is sufficiently close to actuality to be useful. This 
is the case in which the projectile has a conical head. 
The theory determines with exactness the air flow 
past the conical head, and the forces on the head. It 
then can be applied to give estimates of the forces on 
the entire conical-headed projectile. While the results 
do not apply exactly to the ogival-headed projectiles 
used in service, they indicate the way in which the 
moment coefficient (and thus the stability) of an 
actual projectile can be expected to vary with varying 
ogive lengths and velocities. They also pave the way 
for an eventual complete theoretical study of pro¬ 
jectile aerodynamics. 


h This section is based on NDRC Report A-358. 68 


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MOTION OF A SLIGHTLY YAWING CONE 


171 



Figure 7. The conical shock wave set up in air by a 
20-mm projectile in flight, velocity 2,112 fps. (Photo¬ 
graph by courtesy of Ballistic Research Laboratory, 
Aberdeen Proving Ground.) 

The mathematical problem is simplified by noting 
that the air flow in the neighborhood of the conical 
head is the same as if the cone were to extend to in¬ 
finity. This is because at the speeds considered, the 
alterations in the flow that are produced by altering 
the length of the projectile cannot “catch up” to the 
head. 

8 4 2 Nonyaw Theory 

If the conical-headed projectile (which may for the 
present purposes be regarded as an infinite circular 
cone) has zero yaw, the mathematical solution has 
been known for some years. 437,438,495 At supersonic 
speeds, a shock wave travels ahead of the cone. If the 
semiangle of the cone, denoted by 0 8 , is not too large 
(the upper limit being about 56 degrees) and the veloc¬ 
ity U of the cone is somewhat larger than the veloc¬ 
ity «i of sound (the limit depending upon 0 8 , if 0 8 
= 15 degrees, U/a\ should be greater than 1.2), the 
shock wave is also conical, and is attached to the pro¬ 
jectile at the vertex (see Figure 7). The semi-angle of 
the shock-wave cone is denoted by 0 W . 


An important feature of the flow of air relative to 
the projectile is that it is “conical.” This means that, 
besides being symmetric about the axis of revolution 
of the projectile, the air velocity components, (such 
as the pressure and density) are the same at all points 
on any one ray through the vertex. The flow relative 
to the projectile is also steady, since only a short por¬ 
tion of the trajectory is being considered, so that the 
changes in magnitude and direction of U can be dis¬ 
regarded. 

In the papers of Taylor and Maccoll, 437,438 the pres¬ 
sure (and hence the drag) on the conical head, the 
velocity components of the flow, and the shock angle 
0 W , have been determined as functions of 0 S and the 
Mach number U/d\ in excellent agreement with ex¬ 
periment. 

8 4 3 Yaw Theory 

If the projectile in flight yaws through a small 
angle e, the conical shock wave will also yaw, although 
through a different angle 8. It can be proved 68 that it 
is otherwise unaltered, that is, it remains a circular 
cone of semi-angle d w , attached to the vertex of the 
conical head. (See Figure 8.) The flow, as before con¬ 
sidered relative to the projectile, is no longer sym¬ 
metrical but one can still assume that the flow is 
steady, and that it is the same at all points on any one 
ray through the vertex. 

The problem is now one of the steady flow of a com¬ 
pressible ideal gas. The lack of axial symmetry pre¬ 
sents the following complications: (1) there is no veloc¬ 
ity potential, that is, the flow is not “irrotational”; 
(2) the entropy change produced by the shock surface 
varies from streamline to streamline because of the 
ack of symmetry, that is, the flow is not “isentropic”; 



CONFIDENTIAL 









172 


EXTERIOR BALLISTICS OF HYPERVELOCITY PROJECTILES 


and (3) at ballistic speeds the entropy variations and 
the rotation of the flow are small, but they may be¬ 
come appreciable at hypervelocities. 

The method adopted is to regard the flow as differ¬ 
ing from the nonyaw flow only by a small “perturba¬ 
tion.” Products of the perturbation terms are neg¬ 
lected ; this amounts to assuming that e 2 (the square of 
the yaw, measured in radians), can be neglected, or 
that € does not exceed (say) 5 degrees. A number of 
cruder approximate 1 * treatments have been sug¬ 
gested, 47 ’ 434>4451485 ’ 494 some applicable to arbitrary 
slender bodies of revolution; however, they do not a 
priori seem accurate enough for the present purpose. 

Without entering into details, the method can be 
described as follows. The air is considered to move 
with velocity U relative to the stationary cone. Spher¬ 
ical coordinates (r,0,0) are used as indicated in Figure 
9. The velocity components of the air relative to the 



Figure 9. Coordinates used to describe the motion of 
a supersonic cone having a slight yaw e. (This figure has 
appeared as Figure 2 of NDRC Report A-358.) 

conical head in the directions of increasing r, 6, and 0, 
respectively, are denoted by u,v,w. The air density at 
the point (r,0,0) is denoted by p, and the pressure by 
p. The nonyaw values of these quantities are indi¬ 
cated by barred letters u, v, etc.; from symmetry, 
w — 0. The differences, such as u-u, can be expanded 
as Fourier series in 0. It is found that only the “first 
harmonics” are significant, and that one can write 
(for the flow between the shock wave and the conical 
head) 

u = u + ex cos 0, v = v + ey cos 0, w = ez sin 0, 
p — p + erj cos 0, p = p + cos 0, (21) 

where x, y, z, rj, £ are (for given 6 S and Mach number) 
unknown functions of 6 only. 

Five linear differential equations for these five un¬ 
knowns are obtained (neglecting e 2 ) from Newton’s 

1 A closer analysis shows that the results are substantially 

valid if only c 3 is negligible. 111 i The results should thus apply for 

yaws up to 10 degrees. 


equations of motion, the “equation of continuity” 
(expressing the conservation of mass), and the re¬ 
quirement that the flow, once past the shock wave, 
must be adiabatic — that is, p/py (where 7 is the ratio 
of the specific heats) must be constant for each air 
particle, although it varies from particle to particle 
because of the variation in entropy. These equations 
can be greatly simplified since various combinations 
of the unknowns can be integrated exactly; and one 
finally obtains a linear second order differential equa¬ 
tion for x/d as a function of 6, d being an unknown 
constant. When the conditions at the shock wave are 
introduced into the standard Rankine-Hugoniot equa¬ 
tions, the values of x/d and d(x/d)/dd when 6 = 6 W , 
can be derived from the resulting equations. The 
quantity x/d is now determined; and the values of d, 
x, and all the other unknowns are settled by the final 
requirement that the flow must be tangential to the 
conical head. 

8 4 4 Experimental Verification 

A valuable check on the theory is provided by the 
fact that it determines the ratio 8/e of the shock yaw 
to the projectile yaw. Thanks to the cooperation of 
the Ballistic Research Laboratory of Aberdeen Prov¬ 
ing Ground, this ratio was computed in one case 
( d s = 15 degrees, U/ai = 1.901) and compared with 
the actual values as shown by measurements of spark 
photographs (Figure 7) of conical-headed projectiles 
in flight. The results are shown in Figure 10. j Since 
the departures from the theoretical line are all within 
experimental error, the agreement is satisfactory. 

A further check is given by the fact that theoreti¬ 
cally the yaw should be independent of the shock 
angle 0 W . This also holds within experimental error, 
but it is less conclusive since 0 W is rather sensitive to 
variations in Mach number in this range and is hard 
to measure exactly. 

8 4 5 Normal Force and Center of Pressure 

From the theory, also the forces on the conical 
head can be calculated. The force on the head normal 
to the projectile axis, say L H , is expressible as 

L h = pK NH D 2 U 2 , (22) 


i The values given for 8 and e are, strictly, not 8 and e but 
their projections on the plane of the photographic plate. How¬ 
ever, it can be shown that the projections have the same ratio 
as 8 and e, so the comparison still is valid. 


CONFIDENTIAL 

















MOTION OF A SLIGHTLY YAWING CONE 


173 


6*> 

I 

* 

O 

o 

X 

in 



PROJECTILE YAW £ IN DEGREES 

Figure 10. Ratio of shock yaw to projectile yaw, for d s = 15° and U/ai = 1.9. The yaw on each photographic plate is 

represented by O-#, where O and • are the values read by two observers. (This figure is used by courtesy of Ballistic 

Research Laboratory, Aberdeen Proving Ground.) 


where K NH , the contribution of the head to the total 
normal force L, is determined by 

Knh = - y cot e[ux + —(c 2 - U 2 )jp/pt/ 2 , (23) 

evaluated with 6 = 6 S . Here 7 is the ratio of specific 
heats (= 1.405), c 2 = U 2 + 2ai 2 /(y — 1), and p is 
the density of the stationary air. 

This normal force L H acts at a “center of head pres¬ 
sure’’ at a distance %l sec 2 6 S from the vertex, where 
l is the perpendicular distance from the vertex to the 
base of the conical head (see Figure 11 ). 



Figure 11. Idealized conical-headed projectile. (This 
figure is based on Figure 4 of NDRC Report A-358.) 


The axial drag on the head is also determined from 
the present theory. It is, however, the same as in the 
nonyaw theory of Taylor and Maccoll , 437 since the 
change produced in it by the yaw, though significant, 
is proportional to e 2 . Its determination is discussed in 
the next section. 

Estimates can now be made of the normal force, 


center of pressure, and overturning moment for the 
entire proj ectile. For this purpose the simplest assump¬ 
tions are made: namely (a) the projectile has a cylin¬ 
drical body of length c (see Figure 11 ), (b) the pressure 
on its base is zero (which is very nearly the case at 
high velocities), and (c) the perturbation in the pres¬ 
sure produced by the yaw is assumed to vary linearly 
along the cylindrical body. The resulting estimates 
can be improved when more accurate knowledge of 
the air flow along the body of the projectile becomes 
available. It is found that, approximately, 



that the distance from the vertex to the center of 
pressure is approximately 

. D 2 

j = \{2l + c) + 6(7 + c y (25) 

(in which the last term is usually negligible); and, 
finally, that the overturning moment coefficient K M} 
if the projectile is homogenous (in particular, with 
no windshield), is roughly given by equation (26) 

K ' r T Knh (, 12 D + 42>(J + 3c)X 1 + t)' <26) 

Equation (25) for the center of pressure is compared 
in Table 3 with experimental values for conical-headed 
3.3-in. shells , 177 excluding shell types for which the 
values are uncertain. There is agreement to within a 


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174 


EXTERIOR BALLISTICS OF HYPERVELOCITY PROJECTILES 



Figure 12. Normal head force coefficient Knh for projectiles with conical heads having various cone semi-angles, as a 
function of the Mach number U/a\. 


Table 3. Comparison of estimated and experimental 
positions of center of pressure. (3.3-in. conical-headed 
shells: j = distance from vertex of cone to center of 
pressure. All distances measured in calibers.) 


Shell 

type 

l 

L 

Experi¬ 
mental j 

Esti¬ 
mated j 

Discrep¬ 

ancy 

111 

3.26 

2.49 

2.68 

3.03 

+0.35 

112 

3.26 

2.49 

3.05 

3.03 

-0.02 

123 

3.56 

2.49 

3.30 

3.23 

-0.07 

125 

3.26 

1.49 

3.10 

2.71 

-0.39 

168 

2.58 

1.76 

2.25 

2.31 

+0.06 

169 

2.94 

1.76 

2.72 

2.59 

-0.13 


probable error of 0.15 calibers, which is not unreason¬ 
able in view of the experimental errors and the rough¬ 
ness of the estimate leading to equation (25). It should 
be noted that all but the last two shell types in Table 
3 had varying amounts of boat-tail, the effect of 
which is apparently considerable, though it has been 
neglected in making the estimate. 

8 4 6 Values of K NH and 8/e 

The values k of the normal head force coefficient 
K nh and yaw ratio 8/e are shown for various cone 
semi-angles d 8 as functions of the Mach number U/ai 
in Figures 12 and 13. From equation (26), K M should 
vary with the velocity in roughly the same way as 
K nh for a given conical-headed projectile; thus Figure 
12 indicates roughly how the moment coefficient K M 
may be expected to vary with the velocity for super- 

k These values are based on extensive computations on this 
theory that have been carried out recently by the Department 
of Electrical Engineering of Massachusetts Institute of Tech¬ 
nology, under the auspices of the Navy Department. 331 


sonic projectiles in general. It should be remarked 
that there are two values of Knh and 8/e theoretically 
possible for each velocity. This is because the Taylor- 
Maccoll nonyaw theory shows that two regimes are 
theoretically possible in the nonyaw case, having dif¬ 
ferent values of 6 W , and leading to different drags on 
the conical head; however, only one of them is ob¬ 
served in actual firings. The other possibility (not 
shown in Figures 12 and 13) was long believed to re¬ 
present an unstable condition of the projectile, but it 
recently has been observed in carefully controlled 
wind-tunnel experiments. The present theory pro¬ 
vides a partial explanation. It is found that the yaw 
ratio 8/e (and also K N ) is large in the second nonyaw 
regime, so that the yaw always present in projectiles 
as fired would lead to impossibly large shock-wave 
yaws. This shows that the second regime, though 
stable for very small yaws, is unstable for the yaws 
that occur in ballistic practice. 



U/o, 

Figure 13. Ratio of shock-wave yaw 8 to projectile 
yaw e for projectiles with conical heads having various 
semi-cone angles, as a function of the Mach number 
U/a\. See Figure 8 for a representation of the two angles. 


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DRAG COEFFICIENT FOR HIGH-VELOCITY CONE 


175 


As pointed out above, Figure 12 shows in a general 
way how the moment coefficient K M may be expected 
to vary with the velocity for pointed projectiles in 
general although experimental observations are need¬ 
ed to verifjr these relations. Figure 13 reveals the in¬ 
teresting feature that at extremely high speeds the 
shock yaw 8 becomes larger than the projectile yaw e. 
Calculations show that the limiting values of 8 /e (ap¬ 
proached as the speed becomes infinite) are: 1.06 for 
d s = 10 degrees, 1.07 for 0 S = 20 degrees, and 1.08 for 
6 S = 30 degrees. 


8 5 THE DRAG COEFFICIENT FOR A CONE 
MOVING WITH HIGH VELOCITY 

851 Simplification of Nonyaw Theory 

As has been noted in Section 8.4, the theory of Tay¬ 
lor and Maccoll determines the air flow in the neigh¬ 
borhood of the head of a projectile, if the head is 
conical and the yaw is zero. It does not, however, lead 
to explicit formulas for the velocity components 
(u,v) of the flow, or for the drag on the head of the 
projectile, etc.; and such explicit formulas would be 
convenient for both practical and theoretical pur¬ 
poses. A simplified approximate treatment was de¬ 
vised which achieves such explicit formulas. 1 The dif¬ 
ferential equation of Taylor and Maccoll is replaced 
by an approximate equation that can be solved ex¬ 
plicitly. This is done by disregarding the variation in 
air density in the flow behind the shock. In this way 
it is found that 


u/u 8 = sin 2 6 8 + cos 6 8 cos 0 

+ sin 2 0 8 cos 0 In 


tan 1/20 
tan 1/20/ 


(27) 


(where u s denotes the value of u for 0 = 0 S ) with sim¬ 
ilar formulas for v, etc. The drag on the conical head 
can be expressed as 


pK DH D 2 U 2 


(28) 


and K dh is likewise determined approximately. Its 
values, for varying Mach numbers U/ai, are shown 
for two cone semi-angles (0 S = 10, 20 degrees) in 
Figure 14. 

Theoretical estimates of the error produced by this 
approximation show that u and Kdh should be accur¬ 
ate to within 1 percent for Mach numbers of over 2.5 


1 Slight changes in the notation used in NDRC Report 

A-126 25 have been made in the present summary. 



Figure 14. The head drag coefficient Kdh, as a func¬ 
tion of the Mach number U/ai, according to the approx¬ 
imate theory. (This figure has appeared as Figure 3 in 
NDRC Report A-126 where the coefficient was desig¬ 
nated as Ad.) 

and cone angles of 10 degrees or over, and the accur¬ 
acy is much greater for large Mach numbers. How¬ 
ever, the approximations to v and 6 W will be less 
accurate, as confirmed by numerical calculation. 

It is known that, 189,555 roughly, 

K d = 3 K dh , (29) 

where K D is the drag coefficient for the entire pro¬ 
jectile. In this way K D can be estimated, using Figure 
14, for velocities higher than those for which direct 
experimental determinations have been carried out. 

8 5 2 Effect of Yaw 111 

The theoretical evaluation of the extent to which 
the yaw of a conical-headed projectile increases the 
drag on the head is in principle similar to the determi¬ 
nation of the other yaw effects described in Section 
8.4. It is necessary, as shown there, to consider the 
second approximation on the assumption that the 
first approximation is now known. It can be shown 
that the desired second approximation (now neglect¬ 
ing e 3 , where e is the yaw in radians) is expressible in 
the form 

u = u + ex cos 0 + e 2 (w 0 + u 2 cos 2 0), 

v = v + ey cos 0 -f- e 2 (v 0 + ifr cos 2 0), 

w = ez sin 0 4- e 2 w 2 sin 2 0, (30) 

p = p + e ?7 cos 0 + e 2 (po + P 2 cos 2 0), 

p = p + cos 0 + e 2 (po + P 2 cos 2 0), 

where the ten unknowns uo, u 2 , etc., are functions of 0 
alone for given 0 S and U/ai. For the determination of 
the drag, only the unknowns with subscript 0 need to 
be considered. The equations connecting them are de- 


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176 


EXTERIOR BALLISTICS OF HYPERVELOCITY PROJECTILES 


rived and simplified in essentially the same way as 
before (in the first approximation). Two complica¬ 
tions arise, however. First, the equations, though of 
the same general character as before, are decidedly 
more complicated. Second, several of the unknowns 
apparently become infinite when 6 = 0 S . This is be¬ 
cause in these equations e*/v is neglected rather than 
€ 3 , and since v = 0 when 6 = 6 S , the equations are no 
longer valid approximations for values of 6 close to 
0 S . Fortunately the unknown (p 0 ) needed for the effect 
of yaw on the drag is not affected in this way. Finally, 
it is shown that the drag on the head of the projectile 
(i.e., the component of the air resistance along the 
trajectory, rather than along the axis of the projectile) 
is expressible as 

Kdh( 1 + Kdyh£ 2 )pD 2 U 2 , (31) 

where p is the free air density and K DY h, the “head- 
yaw drag coefficient,” is expressed in terms of the 
solutions of two linear second order differential equa¬ 
tions with simple boundary conditions but with com¬ 
plicated coefficients. 111 

8 5 3 Estimate for Entire 

Conical-Headed Projectile 

Using the idealized projectile shown in Figure 11, 
it is assumed (for a rough approximation) that the 
yaw does not appreciably change either the pressure 
on the base (which is nearly zero, in any case, at high 
velocities) or the skin friction. From equation (31),it 
now follows that the total drag on the conical-headed 
projectile is expressible as 

K d { 1 + K D yt 2 )pD 2 U 2 , (32) 

where K DY , the “yaw drag coefficient” (often also de¬ 
noted by Kd 6 2 ) is given roughly by 

r D + (33) 

where c/D is the length of the projectile, exclusive of 
head, in calibers, and rj s and p s denote the values of rj 
and p when 0 = 0 S (known from the first approxima¬ 
tion considered in Section 8.4). 

The computation of K DYH in a typical case ( d s = 15 
degrees, U/ai = 1.954) has been undertaken by the 
Department of Electrical Engineering at Massa¬ 
chusetts Institute of Technology, under the auspices 
of the Bureau of Ordnance, Navy Department. As 
reported, it is known 509 that for standard projec¬ 
tiles K D y = 20. 


8 5 4 Suggestions for Further Research 

Different Head Shapes 

It would be very desirable to have a theoretical 
study of more realistic head shapes than the conical. 
Progress has been made in this direction under the 
auspices of the Armed Forces. One method 332 is based 
on the fact that any pointed ogive can be closely ap¬ 
proximated by a cone in the neighborhood of its tip, 
and the airflow past the ogive is derived as a pertur¬ 
bation from the theoretical flow past the cone. This 
seems to give a good approximation to the flow 
around part, at least, of the ogive. Another proposed 
method 202 is to start with a plausible shape for the 
shock wave (which is hard to predict exactly for 
ogives other than conical) and then determine the 
ogive that would give this shock wave, and the result¬ 
ing airflow. By repeating this computation with sev¬ 
eral shock-wave shapes, the flow for a given ogive 
could be obtained by interpolation. The computa¬ 
tions are necessarily laborious, since now partial dif¬ 
ferential equations must be solved. 

Drag Function 

Once the flow around the head of a (nonyawing) 
projectile has been determined, it is possible, in prin¬ 
ciple, to extend the calculations to give the flow 
around the body of the projectile. If this is practica¬ 
ble, it would enable the drag function of the projec¬ 
tile to be calculated by theory alone, long a dream of 
ballisticians. As mentioned in Section 8.4.3, the ap¬ 
proximate methods previously published do not seem 
to be sufficiently precise. Grave difficulties would 
have to be overcome (particularly in view of the 
boundary-layer effect and turbulence), and a com¬ 
bination of theoretical and experimental work may 
be the most effective. 

Stability Factor 

It should be possible to extend the determination 
of the effects of yaw on the flow, which has been car¬ 
ried out for conical heads, to general ogives or even 
to the entire projectile. The stability factor of a pro¬ 
jectile would then be predictable by theory alone, 
with obvious advantages to the projectile designer. 

Other Aerodynamic Coefficients 

Finally, it may be possible to calculate the effects 
produced by other features of the air resistance (in¬ 
volving the “yawing moment due to yawing,” the 


CONFIDENTIAL 




TRAJECTORY DETERMINATION BY TRACER PHOTOGRAPHY 


177 


Magnus effect, etc.) on the flow past a cone, an 
ogive, or even the complete projectile. A major diffi¬ 
culty here is that the airflow (relative to the projec¬ 
tile) could no longer be regarded as steady, but if this 
problem can be solved, the more recondite aerody¬ 
namic coefficients (which are difficult to measure) 
would be determined, and such things as the drift of 
the projectile and the rate of damping of its yaw, 
would be predictable. 

Effect of Higher Powers of Yaw 

It seems neither desirable nor practicable to repeat 
the method of successive approximations to take into 
account the effects produced by higher powers of the 
yaw. These effects are small in practice; and, more¬ 
over, the method would probably break down on 
account of the infinities introduced by the method of 
approximation. If the way in which K N (say) varies 
with the yaw were desired theoretically, the nonlin¬ 
earity of the fundamental differential equations 
would have to be faced. 

86 TRAJECTORY DETERMINATION BY 
TRACER PHOTOGRAPHY- 

861 General Method 

A simple and convenient method of obtaining bal¬ 
listic data from firings was developed at the Univer¬ 
sity of New Mexico, the original purpose being to 
investigate the relatively unexplored hypervelocity 
region. The method, however, is of general applica¬ 
bility as far as velocity is concerned, but is limited to 
projectiles equipped with tracers. The tracer projec¬ 
tile is photographed in flight at night, using an ordi¬ 
nary (still) camera. A vibrating “chopper” in front of 
the open lens provides an automatic time scale, and 
the resulting photograph gives the complete trajec¬ 
tory and enough data to compute the velocities, re¬ 
tardations, and ballistic coefficient of the projectile. 

8 6 2 Experimental Details 

The setup is illustrated in Figure 15, which is not 
drawn to scale. The projectile is fired from the gun G 
in a vertical plane through the horizontal line MNQ. 
Lights are placed at fixed points M and N, to provide 

“This section is based on NDRC Report A-283, 57 which 
should be consulted for complete details of the method and for 
a discussion of the possible sources of error. 



Figure 15. Setup for trajectory determination by tra¬ 
cer photography. (This figure is a perspective interpre¬ 
tation based on Figures 3 and 4 of NDRC Report A-283.) 


reference points on the photographic image. The 
camera and chopper are set up at C, with the plane of 
the photographic plate perpendicular to a vertical 
plane through the line CN. The plane CQNM is hor¬ 
izontal and CQP is vertical. A third light at L, placed 
near G in the plane CQNM, provides an additional 
reference point and enables the beginning of the tra¬ 
jectory to be located precisely. The chopper is set 
vibrating just before the gun is fired, each period of 
the chopper giving two occlusions of the image. 

The camera lens should have as flat a field as pos¬ 
sible, although it is possible to correct for a moder¬ 
ate amount of distortion. The chopper used was 
simple, consisting essentially of a weighted spring 
anchored to a heavy block; but it worked very well, 
its period (0.0936 sec) varying by less than 0.2 per 
cent with temperature (from — 2 C to +27 C) and 
amplitude. The chopper and camera were mounted 
separately, to avoid transmitting vibrations to the 
camera. With the layout chosen, ranges up to 2,500 
yd were photographed; but ranges several times 
greater could be used with proper choice of lens, chop¬ 
per, and geometry. The errors can be made negligible 
for low-angle fire and high accuracy can be obtained 
by introducing extra lights marking reference points; 
but in high-angle fire it is difficult to correct for lens 
distortion, although this error is not serious with a 
good lens. In designing the layout, it should be borne 
in mind that many tracers do not reach full brilliance 
until nearly 34 sec after leaving the gun. 

8 6 3 Reduction of the Experimental Data 

The coordinates x and y (see Figure 15) of the pos¬ 
ition Pof the projectile corresponding to a point P' of 
the photographic image (see Figure 16) are found to 


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178 


EXTERIOR BALLISTICS OF HYPERVELOCITY PROJECTILES 


P' 



a' b" 1 

*-* - * . L . 

L' M' E' D' 

Figure 16. Measurements on the photographic plate 
in trajectory determination by tracer photography. 
(This figure is based on Figure 5 of NDRC Report 
A-283.) 


be given by 


x = 


nd 


;-k -1 


y = 


a , , 

b' k 1 


(34) 


The distances a, d, k, l, and n are identified in Figure 

15, and the lengths a', b f , and c' are shown in Figure 

16. In that figure L', M' and E' are the images of the 
three lights at L, M, and N which define the baseline. 
The point D' is obtained for successive positions 
along the trajectory by dropping a perpendicular 
from P' to the baseline. The horizontal range R is 
then x + GN, and the vertical height Y of P above 
G is obtained by subtracting from y the height of G 
above the line MN. 

In practice, the positions chosen for P' are the ends 
of the intervals into which the trajectory is divided 
by the chopper. To smooth the results, these points 
are grouped into fours (each four corresponding to a 
complete oscillation of the chopper so that the succes¬ 
sive groups are accurately timed), and for every 
group of four the values of R and Y are averaged and 
taken to correspond to the average time. These av¬ 
erage values of R and F are then plotted, and the 
smooth curve, upon which the points are found to lie 
with remarkable precision, gives the trajectory. (See 
Figure 17.) 


To determine with accuracy the velocity as a func¬ 
tion of the time, some smoothing process must be 
employed. The adopted procedure, which works well 
for low-angle short-range trajectories, is to begin by 
drawing an accurate time-distance curve based on the 
trajectory record of Figure 17, the distance traveled 
being taken equal to R, to a sufficient approximation. 
From this curve, the average velocity over each chop¬ 
per period is calculated. These average velocities are 
plotted against the time, and the resulting points 
should lie close to, but by no means on, a smooth 
curve. A smooth curve is drawn close to these points, 
and is checked by comparing the areas under it with 
the ordinates of the R , £-curve, since 


h 

Udt — R 2 — R\. 

The curve is then adjusted until the agreement is 
sufficiently close, and then it constitutes an accept¬ 
able U, t-mrve. 

The procedure can be repeated to give an r, t-curve 
(where r is the retardation) or an r, U- curve, using the 
fact that 



t 2 — t\. 


In this way the law of resistance for the projectile 
can be established, and its ballistic coefficient can 
be determined. 

This method of determining the resistance function 
has an advantage over the customary chronograph 
methods, since one firing here covers a range of ve¬ 
locities, so that comparatively few firings at different 
muzzle velocities cover the whole velocity range de¬ 
sired. 

It is estimated that, over a 2,000-yd segment of 
trajectory, the position-versus-time data are accurate 
to within less than ± 1 per cent, and the velocities 
are accurate to within +2 per cent. The retardations 


M86 PROJECTILE WITHOUT BALLISTIC CAP PLOTTED POINTS AT 0.0936-SEC INTERVALS 

ROUND 8 FIRED DECEMBER 7 1943 ANGLE OF DEPARTURE 3 3 MILS 

100 
90 
80 
H 70 
w 60 
“■ 50 

* 40 

>- 30 
20 
10 


5 10 15 20 25 30 35 40 45 50 


ROUND 8 FIRED DECEMBER 7 1943 ANGLE OF DEPARTURE 33 MILS 



RANGE R IN I0 2 FEET 

Figure 17. A trajectory determined by tracer photography. (This figure has appeared as Figure 6 of NDRC Report 
A-283.) 


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FIRINGS OF CALIBER .50 PROJECTILES 


179 


are less accurate, because of the necessary differen¬ 
tiations. Longer trajectories could be covered by us¬ 
ing a number of cameras disposed along the trajec¬ 
tory so as to record successive segments. The accur¬ 
acy could be increased at the cost of making the setup 
more elaborate. 

8 7 FIRINGS OF CALIBER .50 PROJECTILES 

To investigate the effects of changes in details of 
the design on the ballistic coefficient and the stability 
factor, an extensive series of firings of caliber .50 bul¬ 
lets was carried out through the cooperation of Aber¬ 
deen Proving Ground in connection with an investi¬ 
gation of pre-engraved projectiles. 11 The results are 
analyzed in detail in Section 27.3. In brief, it may 


n This general investigation was carried out by Division 1 at 
the Franklin Institute. See Chapter 31. 


be noted here that: 

1 . The various changes made in number, position, 
shape and width of rotating bands, in radius of ogive, 
and in boat-tailing (but without greatly affecting the 
overall proportions of the projectiles), altered the 
moment coefficient K M very little, but had significant 
effects upon the drag. 211 Thus, roughly speaking, once 
the proportions and balance of the projectile are such 
that it is adequately stable, other changes in the de¬ 
sign do not greatly change the stability factor, and 
hence can be aimed solely at reducing the drag. 

2. Abrupt leading edges of the rotating band, or 
wider or double bands, and similar features increase 
the drag; a boat-tail reduces the drag, even at hyper¬ 
velocity. 

3. Features, such as double rotating bands, that 
decrease the yaw in the gun are very effective in 
keeping the external yaws small, even if the stability 
factor is decreased. 


CONFIDENTIAL 




Chapter 9 

TERMINAL BALLISTICS OF HYPERYELOCITY PROJECTILES 

By H. S. Roberts a and Walker Bleakney h 


91 DISRUPTION OF A LIQUID BY A 
HYPERVELOCITY PROJECTILE 0 

911 Introduction 

The interest of Division 1 in the problem of the 
disruption of a liquid by a hypervelocity projectile 
was aroused by a British report 426 concerning the ef¬ 
fects produced when a 20-mm projectile, having a 
muzzle velocity of about 4,550 fps, penetrated a self¬ 
sealing gasoline tank, half filled with water. The tank 
exploded with surprising violence^ Since such an ex¬ 
plosion might have considerable tactical importance, 
it seemed desirable to investigate the phenomena. 

912 Firings into Cans of Liquid 

In the initial experiments, several empty tin cans 
about in. in diameter and 3^4 in. high were filled 
with water and penetrated by bullets of different ve¬ 
locities. The ones used were 9.8-g bullets from a cal¬ 
iber .30 Springfield rifle, muzzle velocity about 2,800 
fps, and 2.5-g “half hard” Phosphor bronze bullets 
from a caliber .22 Swift rifle, muzzle velocity about 
4,200 fps. The caliber .30 bullets in some cases opened 

a Deceased; formerly physicist, Geophysical Laboratory, 
Carnegie Institution of Washington. Author of section 9.1. 

b Deputy Chief of Division 2, NDRC, and a Member of 
Division 1. (Present address: Department of Physics, Prince¬ 
ton University.) Author of Section 9.2. 

c The experiments described in this section were performed 
by H. S. Roberts at the Geophysical Laboratory under Con¬ 
tract OEMsr-51 in 1942 and by D. T. MacRoberts and asso¬ 
ciates at the University of New Mexico under Contract 
OEMsr-668 in 1942 as corollaries to broader developments of 
hypervelocity projectiles. They were described in informal 
progress reports submitted by those contractors to Division 1, 
but have not been included in any formal reports. The experi¬ 
ments with caliber .50 bullets described in Section 9.1.2 and 
the shock-wave photographs described in Section 9.1.4 were 
made at the University of New Mexico, the others at the 
Geophysical Laboratory. The mathematical analysis presented 
in Section 9.1.6 is based on two memoranda prepared by 
W. F. G. Swann, Director of the Bartol Research Foundation 
of the Franklin Institute, and submitted on December 29, 
1943, and February 28, 1945, to the Chief of Division 1 under 
Contract OEMsr-533 with the Franklin Institute. 

d A similar phenomenon occurs when a high-velocity projec¬ 
tile penetrates soft animal tissue. It has just been learned that 
this aspect of wound ballistics was studied by the Department 
of Biology of Princeton University for the Division of Surgery 
of the Committee on Medical Research. ( Editor's note.) 


up the seams of the cans but in other cases the seams 
held and little damage was done except for the en¬ 
trance and exit holes and a slight bulging of the sides. 
The caliber .22 bullets, however, invariably burst the 
cans tearing them not only at the seams but often 
elsewhere. 

Since the kinetic energy of the caliber .30 bullets 
was nearly twice that of the caliber .22, it is evident 
that velocity, which was 50 per cent greater in the 
case of the smaller bullets, is a more important factor 
in disruption than is kinetic energy. Some of the cans 
were filled with gasoline in place of water and were 
fired on with the Swift rifle in the expectation that the 
gasoline might ignite. This did not take place and the 
damage was apparently the same as it would have 
been if the cans had been filled with water. 

In a separate series of experiments automobile gas¬ 
oline tanks having a capacity of 16 gal were penetra¬ 
ted by caliber .50 bullets weighing 49 g. Figure 1 is a 
photograph of one of the original tanks. Figure 2 
shows the same tank after being filled with water and 
traversed by the caliber .50 bullet fired with a muzzle 
velocity of 3,500 fps. The tank exploded, that is, the 
disruptive impulse must have been considerably 
greater than was needed to split the tank and it must 
have been applied so suddenly that the inertia of the 
water prevented the first split from relieving the pres¬ 
sure elsewhere. Consequently cracks developed in 
many places and the “follow up” was sufficient to 
tear the tank into several pieces and to distribute 
most of the water over a wide area. 

The experiment was repeated using half the charge 
of powder; the result is shown in Figure 3. This tank 
was torn chiefly at the seams and it remained in one 
piece. In both cases the 2-in.x6-in. plank on which 



Figure 1. Automobile gasoline tank before firing. 


180 


CONFIDENTIAL 



DISRUPTION OF A LIQUID BY A PROJECTILE 


181 



Figure 2. Tank shown in Figure 1 after having been 
filled with water and traversed by a caliber .50 bullet, 
velocity 3,500 fps. 

the tank rested was broken. A third tank, only half 
full of water, exploded when a full-charge bullet pen¬ 
etrated it at a considerable depth below the water 
line; but a fourth tank, also half full, failed to explode 
when the bullet entered at, or just below, the water 
line. 



~ Figure 3. Tank similar to one shown in Figure 2 after 
having been filled with water and traversed by caliber 
.50 bullet fired at half powder charge. 


913 Motion Pictures of Jets 

During Disruption 

Some motion pictures were made of paper tanks 
fired on by the caliber .30 Springfield and by the cal¬ 
iber .22 Swift rifles. The tanks, which were paper 
bags 6 in. square, were filled with water to a depth of 
5}/2 in. so that the volume of water was approximately 
a 6-in. cube. The bullet entered horizontally about 
the center of one face from a direction approximately 
parallel to the adjacent faces. The tank stood on a 
wooden box placed near the center of a heavy plank 
spanning two logs. The camera was set up about 90 
degrees from the line of fire and took 16 pictures per 
second. 

Figure 4 shows the effect of a 2.5 g caliber .22 Phos¬ 
phor bronze bullet, velocity about 4,200 fps. The gun 
was outside the picture to the left. The half frame at 
the top shows the tank on the wooden box just before 
firing. The other three frames show most of the water 
being expelled in five jets, approximately normal to 
each of the five free faces of the cube. In the last two 
frames the vertical jet has reached the top of the pic¬ 
ture, about 11 ft above the top of the tank. The pat¬ 
tern on the ground showed that water had been 
thrown about 20 ft toward the gun and about 15 ft 
to the sides. These lobes were nowhere more than 4 ft 
wide and there was no water in the 45-degree posi¬ 
tions except quite close to the tank. A downward im¬ 
pulse on the bottom of the tank was shown by the 
fact that the resilience of the plank caused the wooden 
box to be thrown a foot or so upward. 

The effect of a standard 9.8-g caliber .30 bullet, 
velocity about 2,800 fps, is shown in Figure 5. In the 
half frame at the top the bullet evidently struck to¬ 
ward the end of the exposure. There is a jet about 4 ft 
long directed away from the gun and some evidence 
of a jet starting toward the camera; but the vertical 
jet has not yet appeared. In the last frame the verti¬ 
cal jet has reached its maximum height of only 4 or 

5 ft. 

914 Shock Waves in 
Water Penetrated by a Bullet 

Spark photographs were made of caliber .22 lead 
bullets, velocity 1,200-1,400 fps, passing through 
water in various glass containers. In every case the 
bullet entered through a paper closure at the mouth 
of the container and not through the glass. In Figure 

6 the bullet has entered at the left and has traveled 


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182 


TERMINAL BALLISTICS OF HYPERVELOCITY PROJECTILES 



Figure 4. Motion picture of a paper tank penetrated Figure 5. Motion picture of a paper tank penetrated 

by a caliber .22 bullet, velocity 4,200 fps. by a caliber .30 bullet, velocity 2,800 fps. 


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DISRUPTION OF A LIQUID BY A PROJECTILE 


183 



two thirds of the way to the end of the jar. The shock 
wave that precedes the bullet has cracked the glass in 
many places but there is no evidence that the frag¬ 
ments have moved, except that thin sheets of water 
are seen emerging from the cracks at the top of the 
picture. 

Following the bullet is a “conical” disturbance 
whose nature has not been determined. Its velocity 
of propagation in the water was estimated to be 
about 66 fps. The sharply defined reflecting surface 
of this disturbance was interpreted to indicate the 


Figure 6. Spark photograph of a glass jar filled with 
water being traversed by a caliber .22 bullet, velocity 
1,200-1,400 fps. 


Figure 7. Spark photograph of a spherical glass tank 
not quite full of water being traversed by a caliber .22 
bullet, velocity 1,200-1,400 fps. 


presence of either a definite liquid-vapor interface or 
else an extremely fine emulsion. 

That these phenomena do not depend on the shape 
of the container was shown by firings into other ves¬ 
sels. Figure 7 shows a spherical container not quite 
full of water in which the bullet has nearly reached 
the far wall. The glass is cracked much as it was in 
Figure 6, the fragments remain in place and the water 
level has not been disturbed. In Figure 8A taken just 
before firing, we have a large glass tube filled with 
water except for a small bubble near the right hand 
end. Figure 8B shows the bullet about half way 
through. The conical disturbance has reached the 
wall of the container, and glass and water are moving 
outward with a velocity much greater than that at 
which the conical disturbance was propagated. The 
bubble has moved a few millimeters away from the 
bullet. 

Previously published photographs 486 of shock 
waves in water penetrated by a bullet showed that 
the shock wave will not damage a flexible tank wall 
but will damage a rigid one. A curved shock wave was 
reflected from a cellophane wall of a small tank with¬ 
out damage to the cellophane. 


Figure 8A. Spark photograph of a large glass tube 
filled with water; before firing. 


Figure 8B. Same tube as in Figure 8A; caliber .22 bul¬ 
let part way through, velocity 1,200-1,400 fps. 


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184 


TERMINAL BALLISTICS OF HYPERVELOCIT Y PROJECTILES 


915 Transient Pressures in 

Water Penetrated by a Bullet 

Measure of the Bursting Strength of a Tank 

These purely qualitative experiments have shown 

(1) that the disruptive effect increases rapidly with 
the velocity of the bullet and that the velocity is a 
more important factor than the kinetic energy; and 

(2) that there are two phenomena, (a) a shock wave 
which originates when the bullet enters the liquid and 
travels with the velocity of sound (about 4,800 fps in 
water), and (b) a secondary disturbance of unknown 
nature which travels much more slowly. They show 
further that much more energy is associated with the 
secondary disturbance than with the shock wave, at 
least at low bullet velocities. 

The ability of the container to resist disruption by 
a uniform pressure (such as might be applied by a 
pump) can be expressed either (1) as the pressure just 
before the tank bursts, or (2) as the work that must 
be done to stretch the tank to the bursting point. The 
latter is the better concept for our purpose because 
we have, in the shock wave, pressures that may be 
far greater than the bursting pressure of the tank but 
of such short duration that the work they can do may 
be inadequate. 

We can imagine two tanks of the same size, one of 
glass and the other of rubber, with the wall thickness 
so proportioned that they will burst at the same pres¬ 
sure. Once the tank is filled, a single stroke of the 
pump might be sufficient to burst the glass tank while 
many strokes might be required for the rubber tank. 
As we have seen, the pressures we are dealing with 
are nonuniform. Even so, the work done on the walls, 
or on one wall, of a particular container should be a 
measure of the relative effectiveness of the same type 
of bullet fired at different velocities; although there 
is the possibility that this method may fail to indicate 
significant changes in pressure distribution. 

Experiments with Water Ballistic Pendulum 

The Pendulum. These considerations led to the con¬ 
struction of a water ballistic pendulum, shown in 
Figure 9. The body, D, of the pendulum was a 16-in. 
length of a worn 3-in. steel gun liner having walls 
about 1-in. thick. The back (right hand end in the 
figure) was closed by a heavy steel plate P, held on 
by eight J/£-in. cap screws and made watertight by a 
rubber gasket. The front of the body was closed by a 
disk of tar paper held in place by the flanged nozzle N. 



The purpose of the nozzle was to confine the expelled 
water to a narrow cone so that the momentum of the 
empty pendulum would be substantially equal to the 
total momentum of the water. 

The pendulum was hung from the ceiling by four 
strap iron links L and its angular deflection was indi¬ 
cated by a friction pointer which Avas moved over a 
scale S by a pin in an extension of one of the links. A 
piezoelectric pressure gauge G fitted into a hole in the 
rear plate Avas connected to an oscillograph to give a 
permanent record of the water pressures. The pendu¬ 
lum body could be filled completely with water by 
disengaging the rear links so that the body hung 
vertically. 

Procedure. The bullet was fired through the tar 
paper disk as nearly as possible along the axis of the 
body. Practically all of the water A\ T as expelled vio¬ 
lently, taking the bullet with it. Thus, if we assume 
that no energy was lost in stretching the pendulum, 
(or the cap screAv) the kinetic energy, e p , imparted to 
it is given by equation (1) 

e v = Wr{\ — cos 0) foot-pounds (1) 

where W is the Aveight of the empty pendulum plus 
two of the links (71 lb); r is the radius of the links 
(5.65 ft) and 0 is the angle of deflection. 

Very short bullets Aveighing only 1.8 g were made 
of “half hard” Phosphor bronze, and by modifying 
the poAvder charge a range of velocities up to 5,800 
fps was obtained. Only 2 rounds Avere fired at this 
highest velocity because the poAA^der pressures Avere 
found to have been dangerously high. Bullet velocity 
was determined for each round by means of two 
screens of 0.003-in. copper wire placed 8 ft apart; the 
velocities are believed to be accurate to about 2 per 
cent 

Measurements of Pressure and Kinetic Energy. The 


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DISRUPTION OF A LIQUID BY A PROJECTILE 


185 



Figure 10. Kinetic energy of water ballistic pendulum 
(full curve) and 0.0145 times the kinetic energy of the 
bullet (dashed curve), as functions of the bullet velocity. 


pressure gauge was designed originally for recording 
powder gas pressures; its natural frequency, about 
70,000 c was too low to permit adequate represen¬ 
tation of the shock waves. Thus its indications may 


A 


\ 



Figure 11 A. Pressure oscillogram, water ballistic pen¬ 
dulum; bullet velocity 2,580 fps, first pressure pulse 
10,800 psi. 


have been considerably lower than the actual peak 
pressures. They should, however, provide a basis for 
comparison. In Figure 10 the kinetic energy imparted 
to the pendulum is plotted against the velocity of the 
bullet. The ordinates of the dashed curve are propor¬ 
tional to the kinetic energy of the bullet. Below 2,000 
fps the efficiency (e p /eb) is quite low after which it 
rises to a flat maximum at about 3,500 fps and then 
falls off. There seems to be no reason to expect a 
maximum efficiency to occur in general at 3,500 fps. 
Its position may be due to the shape, dimensions, or 
rigidity of the pendulum and may be very different for 
an actual gasoline tank. 

In Figure 11 are two pressure oscillograms for the 
ballistic pendulum. They were made with a still cam¬ 
era, the oscilloscope operating with a 500-c recurrent 
sweep; the records are thus 2 msec long, time being 
read from left to right. Since the camera shutter was 
open for 1/25 sec the base line was swept over many 
times before and after the shot and the heavy fogging 
covers up the record of the lower pressures. For A the 
bullet velocity was 2,580 fps and the peak pressure of 
the first pulse to arrive was 10,800 psi. In B the ve¬ 
locity was 5,800 fps and the peak pressure of the first 
pulse 78,000 psi. The presence of a series of shock 
waves can best be explained by multiple reflections 
from the wall of the tube. In Figure 12 the peak pres¬ 
sure of the first shock wave is plotted against bullet 
velocity for several experiments. The pressure begins 
to increase rapidly at about 2,000 fps and appears to 
go through a definite maximum at the velocity of 
sound, 4,800 fps. 


B 





Figure 11B. Pressure oscillogram, water ballistic pen¬ 
dulum; bullet velocity 5,800 fps, first pressure pulse, 
78,000 psi. 


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186 


TERMINAL BALLISTICS OF HYPERVELOCITY PROJECTILES 


— 


O 

8 



o 

° o 

o 

o 

o 

c 

0 U- 

o 

o o 

5 O 



°0 2000 4000 6000 

BULLET VELOCITY IN FPS 


Figure 12. Shock-wave pressure as a function of bullet 
velocity, water ballistic pendulum. 



psi for nearly a millisecond; and then there is an¬ 
other short period of somewhat lower pressure begin¬ 
ning near the end of the sweep and continuing over 
into the early portion of the following sweep. 


Experiments with a Heavy Steel Tank 

Before the water ballistic pendulum was built some 
work was done with a rectangular steel tank, Figure 
13. The front sides, top and bottom were 34-in • thick 
and the back 1-in. thick. The 34 _ in- plates were 
welded to each other but were bolted to the back 
plate. Four circular windows were provided to fire 
through and were closed by disks of tar paper. 
There were two holes in the back plate for the 
pressure gauge; the one used was 5% in. from the 
line of fire. A considerable number of experiments 
were planned for this tank, but the first shot opened 
up the weld at the bottom of the front face. The tank 
was kept in operation for a while by calking the leak 
with putty, but when the side seams began to open up 
its use was abandoned. 

Figure 14 is a pressure oscillogram for a 3.1-g soft 
nose bullet fired from the caliber .22 Swift with a ve¬ 
locity of 3,000 fps. The pressure reading for the initial 
shock wave is 7,800 psi. There is a rather violent os¬ 
cillation of the gauge, or of the back plate, but the 
average pressure remains close to zero until a second 
shock wave arrives about 0.4 msec after the first. The 
average pressure remains in the neighborhood of 1,000 


Conclusions from the Pressure and Energy 
Measurements 

The early deductions from the qualitative experi¬ 
ments were, on the whole, confirmed by the records 
of energy and pressure. The work done on the water 
ballistic pendulum increases about sevenfold as the 
bullet velocity is raised from 2,000 to 4,000 fps; 
above 4,000 fps the increase is less rapid. The pres¬ 
sure of the initial shock wave may be very high, but 
the pulse is of such short duration that it has little 
energy. The shock wave is followed by an interval 
during which the pressure is close to zero; then the 
pressure rises to a moderate value which is sustained 
for some time. It is believed that disruption of a flex¬ 
ible tank is caused by this later, sustained pulse. The 
mechanism by which this pulse is produced was not 
determined. 

It is quite possible that the water ballistic pendu¬ 
lum may be too unlike the self-sealing gasoline tanks 
we are interested in. A better representation might be 
given by a paper tank in which the surface of the 
water was covered with floating cubes of wood rough¬ 
ly corresponding to an upper, horizontal wall of the 
tank. The height to which one of these blocks is 


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DISRUPTION OF A LIQUID BY A PROJECTILE 


187 


thrown, multiplied by the weight of the block, would 
be somewhat less (due to air resistance) than the 
work done on the base of that particular block. 

Recommendations. In a continuation of this inves¬ 
tigation it would be desirable to study the disruptive 
effects of different types and weights of bullet fired 
from the same gun, modifying the powder charge, if 
necessary, because of changes in weight. The means 
for evaluating the disruptive effect may be rather 
crude (the wood blocks already referred to should be 
adequate) because the effect is likely to vary consid¬ 
erably from round to round. It should be remembered 
that in combat the bullet has to travel for consider¬ 
able distance through air and may have to pierce the 
armored skin of the airplane before it reaches the gas¬ 
oline tank. 

916 Partial Analysis of the Phenomena 

Before these experiments were begun it was known 
that the impact of a high velocity bullet on a body of 
liquid would produce a shock wave in the liquid and 
that if the striking velocity of the bullet were in¬ 
creased, the intensity of the shock wave would be 
increased, probably reaching a maximum intensity 
when the striking velocity equals the propagation 
velocity of the shock wave. The following analysis 
has to do with the mechanism by which this initial 
shock wave, presumably, is produced. Since the anal¬ 
ysis takes no account of the velocity with which the 
shock wave is propagated it does not predict the max¬ 
imum observed in the data of Figure 12 at the veloc¬ 
ity of sound. Later, when the experiments began to 
indicate that the initial shock wave may have very 
little to do with the disruption of a flexible tank, an 
independent analysis of the disturbance set up in the 
liquid as the bullet travels through it was attempted. 
This later analysis was foredoomed to failure because 
the nature of the disturbance was not, and is not, 
known. It seems worth while however to include the 
first analysis even though the initial shock wave does 
not appear to be directly relevant to the subject at 
hand. 

Generation of the Initial Shock Wave 

For the initial shock wave, the fundamental con¬ 
sideration is the speed with which the projectile 
makes room for itself in the container by compressing 
the liquid. If it were inserted very slowly, the ulti¬ 
mate pressure created would be uniform throughout 


the liquid. It would amount to {w/W)E, where w is 
the volume of the projectile, W the volume of the 
liquid, and E the bulk modulus of the liquid. For 
water E is taken as 2 X 10 10 dynes per square cen¬ 
timeter. Thus if a caliber .22 bullet (volume about 
cm 3 ) entered a liter of water slowly, the contraction 
would be about 1 part in 2,000 and the pressure re¬ 
quired would be about 10 atmospheres. 

When the bullet enters the liquid in a time short 
compared with the time necessary for elastic waves to 
reach all parts of the surface of the liquid, the pres¬ 
sure is locally much larger than for the case of slow 
compression. This is true because the amount of liq¬ 
uid that suffers the compression is much less, being 
in fact, merely that liquid in the immediate vicinity 
of the bullet through which the sound wave has 
passed during the time of penetration of the bullet. 

In order to arrive at an approximate expression for 
the pressure in the liquid we may consider the sim¬ 
plified situation consisting of a bullet of radius 
a in the center of a container of radius b. Suppose 
that the radius of the bullet suddenly increases 
at a rate v cm per sec. The increase in size of the 
bullet need last only an instant in order to generate a 
pressure wave, although the longer it lasts the greater 
will be the extent of the wave. The radial displace¬ 
ment $ of a particle of the liquid caused by this acous¬ 
tical wave obeys the differential equation 

i_ = n 2_ ar f2 v 

c 2 dt 2 dr 2 r dr’ 

where r is the distance of the particle from the center 


I 



Figure 14. Pressure oscillogram for water in a steel 
tank fired into by a caliber .22 bullet at 3,000 fps; first 
pressure pulse: 7,800 psi. 


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188 


TERMINAL BALLISTICS OF HYPERVELOCITY PROJECTILES 


of the container, c is the velocity of an acoustical 
wave in the liquid, and t is the time. 

The general solution of equation (2) is given by 
equation (3), 

f = yf(ct - r) + yF(ct + r), (3) 

in which / and F are arbitrary functions. For out¬ 
going waves only / applies, so that equation (3) may 
be replaced by equation (4) 

t = yf(ct - r). ( 4 ) 

It is well known that for such a system the pressure 
p at any point is given by equation (5), 



where E is the elastic modulus of the liquid. By com¬ 
bining equations (4) and (5), the pressure may be 
expressed by equation (6), where p is the density. 



In the case of the spherical bullet that suddenly 
begins to increase in radius at the rate v, the pressure 
at a radius a is given by equation (7). 

P a = pcv. (7) 

In order to obtain the pressure at a later time t at a 
radius r, equation (4) may be differentiated with re¬ 
spect to time and then evaluated for t = 0 to give the 
velocity of particles at radius a and for t = ( r—a)/c 
to give the velocity of the particles at radius r. The 
resulting expressions are given by equations (8) and 



which, when combined with equations (6) and (7) 
give equation (10) for the pressure at radius r. 



The pressure at the wall of the container of radius 
6, when the pulse reaches it, is given by equation (11). 



If d£/dt is of the order of magnitude of the velocity of 
the bullet, say 3,000 fps, and if a/b = 1/15 (a ratio 
that is valid for the cans used in the experiments de¬ 
scribed in Section 9.1.2), p b would be about 6 tons per 
square inch. If a rigid container wall is perpendicular 
to the plane of the wave front, the wave will be re¬ 
flected back along the path by which it came, and 
during reflection the incident and reflected waves will 
combine to give a pressure twice that just stated. 

Equation (11) indicates that for given values of b 
and v the pressure at the wall of the container is pro¬ 
portional to a, and hence to the cube root of the mass 
of the bullet. Equation (11), however, does not show 
why the hypervelocity caliber .22 bullets fired at a 
velocity of 4,000 fps were so much more damaging 
than the caliber .30 ones fired at 2,700 fps. The cal¬ 
culated pressure in the former case was only 9 per 
cent greater than in the latter. 

Transfer of Momenta 

Another feature of the experiments described in 
the foregoing sections is that when a bullet is fired 
into water there is a*very large transfer of momenta 
if the water is free to move. Thus in the case of the 
experiments with the ballistic pendulum the momen¬ 
tum Pi imparted to the pendulum in the direction of 
flight of the bullet was from 13 to 19 times the orig¬ 
inal momentum of the bullet. Because of the law of 
conservation of momentum, the amount of momen¬ 
tum P 2 imparted to the water in the opposite direc¬ 
tion must have been equal to Pi less the original 
momentum of the bullet. 

Such a large transfer of momentum, although it is 
not frequently encountered, is merely a special case 
of a general relationship that may be expressed in the 
following terms. Suppose that a mass m having a ve¬ 
locity v is permitted to share its momentum with two 
other masses. For simplicity in computation suppose 
further that these two masses are equal, and hence 
may be designated by M. For the sake of generality, 
suppose that only a fraction a of the original kinetic 
energy of the first body is conserved as observable 
kinetic energy, the remainder being converted into 
heat; and further, that of the kinetic energy lost by 
this body, only a fraction (3 is conserved as observable 
kinetic energy in the two masses M, the remainder 
being converted into heat. 


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ARMOR PERFORATION 


189 


These two conditions are expressed by equation 

(12) and the conservation of momentum by equation 

(13) . 

■jMVf + yikflV = A /3 (1 - a)mv* (12) 

MV i + MV 2 = «ms • (13) 

Equations (12) and (13) may be solved for the mo¬ 
menta of the two large masses, as given in equations 

(14) ,in which M' represents [2 /? ( l—a)Mm — a 2 m 2 ] 1/2 

I* r T r amv . M'v x 

MVi = (14a) 

MV. = -y (14b) 

Thus by making M sufficiently large, we can make 
the momentum MVi as large as we please; but by 
whatever amount it exceeds amv/ 2, there is always a 
corresponding amount of negative momentum in the 
other mass. It should be remembered that by putting 
a finite amount of momentum into a very large mass, 
we endow it with only an infinitesimal amount of en¬ 
ergy, whereas by putting a finite amount of energy 
into a very large mass, we endow it with a very large 
momentum, an infinite amount in the limiting case of 
an infinite mass. 

92 ARMOR PERFORATION® 

921 Specific Limit Energy 

It is convenient to discuss the perforation of armor 
plate in terms of the “specific limit energy/’ which is 
defined by the expression WVi 2 /d 3 , where W is 
weight of projectile; d is diameter of projectile, that 
is, the caliber; and Vi is limit velocity. By “limit 
velocity” is meant the minimum velocity required to 
defeat f the plate. Most experimental results may be 


e Only a short description is given here of the perforation of 
armor by hypervelocity projectiles since the subject is covered 
much more extensively in the Summary Technical Report of 
Division 2, where a bibliography will be found. 

f The term “defeat” as used here may be defined in various 
ways. The limit velocity for complete perforation of the plate 
by the entire projectile constitutes defeat in Navy terminol¬ 
ogy, whereas a pinhole made by the very tip of the missile 
represents defeat in Army parlance. Limit velocities for bulg¬ 
ing, spalling, cracking, plugging, may likewise be defined. The 
discussion given here applies regardless of which kind of “de¬ 
feat” is under consideration but quantitative and numerical 
results refer to the complete perforation of the plate by the 
armor-piercing core of the projectile. 


described for practical purposes by formulas express¬ 
ing WVi 2 /d 3 as a function of e/d and 0, where e is 
plate thickness, e/d is plate thickness in calibers, and 
6 is obliquity, that is, the angle between the trajec¬ 
tory and the normal to the plate. 

One way to systematize the observations is to 
make a chart in which WVi 2 /d s is plotted against e/d. 
The results for a particular combination of projectile 
and plate material and a particular angle 0 will be 
found to lie along a curve or band, the width of the 
band representing the uncontrollable scatter in the 
data. For a different angle 0 a different band will be 
found. The advantage of such a choice of variables is 
that it reduces the results obtained with all sizes of 
projectiles to a common basis. The fact that exten¬ 
sive investigation has shown that this procedure is 
possible without much error means that there is very 
little “scale effect” in armor perforation. However, 
this scale effect, while small, is nevertheless real and 
is in the direction of decreasing WVi 2 /d z with increas¬ 
ing d. A chart such as that discussed in the preceding 
paragraphs is presented in Figure 15 where the shad¬ 
ed bands indicate the actual performance attained 
with practical projectiles having armor-piercing cores 
of tungsten carbide. It will be noted that the weight 
and diameter refers to this core only, without regard 
to any jacket or cap material. 

A very carefully controlled series of laboratory ex¬ 
periments 158 with nonshattering shot at normal inci¬ 
dence is represented by equation (15). The units are 
the same as those in Figure 15. 

WVi 2 / e V - 26 

~dT = 15Sxl0 \l) ^ 

For the caliber (0.244 in.) and the armor used, this 
represents something very close to the ultimate in 
projectile performance and therefore represents a 
boundary beyond which we do not expect to go with 
conventional projectiles. The boundaries for other 
obliquities have not as yet been so clearly defined. It 
may appear that some limit energies obtained from 
Figure 15 are less than those given by the above rela¬ 
tion. This discrepancy may be related to two circum¬ 
stances, (1) the small scale effect which favors the 
larger projectiles and (2) the fact that the data plot¬ 
ted in Figure 15 refer to the core only whereas the 
carrier in many cases contributes to the penetration 
by giving the core a boost from behind when it strikes 
the plate. 

It has been found that most observations on armor 
perforation in which the projectiles are but little de- 


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190 


TERMINAL BALLISTICS OF HYPERVELOCITY PROJECTILES 


V 

2300 

2400 

2500 

2600 

2700 

2800 

2900 

3000 

3100 

3200 

3300 

3400 

3500 

3600 

3700 

3800 

3900 

4000 

4100 

4200 

4300 

4400 

4500 


LIMIT VELOCITY 

feet/second. 


2(W/d 5 ) V^; lb/m 3 (ft/sec) 2 



British 
2 pdr S.V.Mkll J 

(steel cap) 


W * height of core, pounds 
d * dimeter of core, inches 
V/* perforation limit vel¬ 
ocity feet/second 

e = thickness of plate,inches 


55 MO* 

50 











JjN 

UNCAF 

s, 

*PED 


45 












rCAPPE 

:d 

40 











■B 

Hr 


55 




A 

m 

SEl 

AJi 

[iE 


I 

■P 

w 



do 













ipi 


25 










IP 

py 





20 








Is 1 







15 







. 


W* 












| 




SA 

FE 




5 















O 
















7 8 

e/d = PLATE THICKNESS 

calibers 


DIAMETER OF CORE 

inches 



American 76mm M93 
British I7pdr D. S.- 

British 6pdr D. S. - 
British 2pdr S.V.I'fcll (steel cap) - 
German 50mm AP40 - 


987654321 

e = PLATE THICKNESS 

inches 

The chart shows the way in which limit velocity for complete perforation is related to thickness of homogeneous armor 
(BHN 220-330) perforated by capped and uncapped tungsten carbide cored projectiles striking at normal incidence and at 
30° to the normal. The data represent projectiles having cores ranging from 0.65 to 1.52 inches in diameter, indicated 
on the nomogram scales are projectiles with cores standardized for field use. 


Because of inherent scatter of firing data, results are presented as bands. For each obliquity, the band was drawn to 
include 90% of the points. Capped and uncapped projectiles at 0° scatter randomly through the'band; at 30° there is an 
evident separation as indicated. Tungsten carbide cores usually break up on impact. If there is complete disintegration 
of the core, perforation of the indicated thickness may not be attained. 


EXAMPLE A 

Given an uncapped American 76mm M93 
projectile striking at 30°witha vel¬ 
ocity of 3200 ft/sec, two values of 
plate thickness, about 5 inches and 
inches, are read by following the 
line to each of the two borders of 
the band. It is reasonable to assume 
that plate thicknesses greater than 
frg inches will be safe against this 
projectile, while thicknesses less 
than 5 inches will be vul-nerable. 



EXAMPLEB 

Similarly, forthe same projectile (see 
Example Ajfiredat a plate about 5 in. 
thick, two values.of the striking velo¬ 
city are read by following the line to 
each of the two borders of the band, 
namely 3200 and 2770 ft/sec. The plate 
referred to is likely to be safe againstj 
striking velocities less than 2770 ft/sec 
and vulnerable to velocities greater 
than 3200 ft/sec. 



SOURCE: Based on experiments by Anerican, British and Canadian military establishments. 

Figure 15. Perforation of homogeneous armor by tungsten carbide cored projectiles. (From OSRD Report No. 6053 
Weapon data: fire, impact , explosion , Division 2, NDRC.) 


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ARMOR PERFORATION 


191 


formed in the process can be represented fairly well 
at obliquities near the normal by a relation of the 
form of equation (16), 

WVi 2 (e\* nr , 

~dT ~ R Kdb (16) 

where R is determined chiefly by the strength of the 
plate material and n has a value between 1 and 2. If 
n is given the value 1.5 the De Marre formula in use 
by the Army Ordnance Department is the result. The 
value n = 1 gives essentially the Thompson formula 
used extensively in the Navy Bureau of Ordnance. 
Neither of these forms fits the observations over a 
very wide range without changing R. In fact, for a 
given projectile and plate material, the behavior can¬ 
not be represented over extreme ranges of Vj and e/d 
by any one set of values for R and n. Equation (15) 
has the form of equation (16) and n lies between the 
Thompson and De Marre values. 

The preceding discussion has been concerned with 
projectiles which suffered little deformation on im¬ 
pact with armor plate or in other words “successful” 
projectiles. If the shape of the missile, or more par¬ 
ticularly, its armor-piercing core, is seriously changed 
on impact, the penetration is drastically reduced. 
There is a common belief in ordnance circles that at 
very high velocities, deformation of the bullet is un¬ 
important. Experiment has shown that this idea in 
general is false. The greatest limitation by far on the 
use of hypervelocity in armor penetration is the diffi¬ 
culty in avoiding the shatter of the projectile. Con¬ 
siderable progress has been made in this direction by 
the addition of caps, pads, and other parts which 
aid in introducing the core into the armor without 
shatter. 

The value of an armor-piercing cap is illustrated 
by Figure 16, which shows a caliber .80 steel projec¬ 
tile and the hole it made in a piece of homogeneous 
armor (Brinell hardness: 250) 2.5 in. thick, which it 
had struck at a velocity of about 4,800 fps at an angle 
of 30 degrees from the normal. The projectile had 
been equipped with an armor-piercing cap that dis¬ 
integrated in the process. A similar steel projectile 
without the cap was severely deformed by the impact 
and did not perforate the plate. 

This example illustrates the fact that steel shot, if 
properly designed, stands up against this type of tar¬ 
get under hypervelocity conditions. Steel suffers in 
comparison with tungsten carbide in having lower 
density and lower hardness but may be preferred in 
some hypervelocity applications because of its avail- 



Figure 16. Perforation of homogeneous armor 3.1 cal¬ 
ibers thick by a caliber .80 steel projectile at 30° obli¬ 
quity. (Photograph by courtesy of Princeton University 
Experiment Station, Division 2, NDRC.) 

ability and ease of fabrication. Experience indicates 
that tungsten carbide can be made to outperform 
steel at hyper velocities but nevertheless steel has 
given a creditable performance in experimental fir¬ 
ings under these conditions. It is not yet certain that 
steel projectiles made by customary mass-production 
procedures will do so. 

The design of practical projectiles to meet condi¬ 
tions of impact at high obliquities, spaced armor, face- 
hardened armor and combinations of these factors is 
very complicated and largely empirical at the present 
time. More extensive discussion of these problems is 
beyond the scope of this short review.® 

9 2 2 Armor Perforation by 

Hypervelocity Projectiles 

It is impossible to separate the fields of interior, 
exterior, and terminal ballistics; solve their problems 
separately; and obtain the best overall answer to the 
ballistic problem. Nevertheless it is helpful to make 
an approximate separation in order better to organize 
our thinking. If perforation of armor is the only con¬ 
sideration and the discussion is based on equal ener¬ 
gies of the projectiles when striking the plate, the 
following conclusions may be made as a result of in¬ 
vestigation in the hypervelocity field. 

1. A nondeformable 'projectile gives maximum per¬ 
formance. Failing to meet this requirement in its en- 


* They are treated in Chapter 6 of the Summary Technical 
Report of Division 2, NDRC. 


CONFIDENTIAL 









192 


TERMINAL BALLISTICS OF HYPERVELOCITY PROJECTILES 


tirety the statement still applies to the armor-pierc¬ 
ing core. If concession must be made to deformability 
the amount of deformable material should be held to 
a minimum. There may be exceptions to these state¬ 
ments under special circumstances, such as impact on 
thin armor at high obliquity. 

2. It is advantageous to concentrate the energy in a 
packet of small dimensions. It is easier to punch a 
small hole through a given plate than it is to punch a 
large one. However, a needle is not practicable be¬ 
cause of the difficulty in satisfying condition (1). It is 
desirable, therefore, to make the armor piercing projec¬ 
tile of material having the highest possible density 
and the highest possible strength. The best material 
so far developed for the purpose is tungsten carbide. 
To be sure, this material usually suffers some frac¬ 
tures, but if properly designed, its penetration prop¬ 
erties and destructive power on emerging from the 
plate are such as to make it a formidable weapon. 

3. If two identical projectiles are fired with different 
energies the faster one will pierce the thicker armor at 
zero range and its advantage over the other will increase 
with range. The indications are that if two similar 
projectiles of different caliber are fired with the same 


energy, the smaller will pierce more armor and its ad¬ 
vantage will (percentagewise) increase with range. 
This important conclusion must be suitably qualified. 
The energies must be below the shatter point for the 
projectile but above the velocity of sound. The sweep¬ 
ing generalization that one sometimes hears, “very 
high velocities are unprofitable because the rate of loss 
in velocity is so high,” is false from this point of view. 

4. Muzzle velocities above the critical shatter velocity 
will improve the armor piercing capabilities beyond the 
range where the velocity drops below the shatter point. It 
has been amply demonstrated that a projectile will 
shatter under certain conditions if the striking ve¬ 
locity is above a critical value and fail to pierce the 
plate, whereas at a lower velocity perforation will be 
achieved. 

All of these conclusions point to the fact that for a 
conventional gun as it is used at present the armor¬ 
piercing performance can be improved by adopting a 
tungsten carbide subcaliber projectile with a carrier 
of minimum weight. A successful design of such a 
projectile mounted in a sabot is described in Chapter 
29 and a design of one mounted in a projectile for a 
tapered bore gun is described in Chapter 30. 


CONFIDENTIAL 



PART III 


GUN EROSION 



Attempt the end, and never stand to doubt 
Nothing’s so hard but search will find it out 
—Robert Herrick 
“Seek and Find” 


CONFIDENTIAL 

















































































































































































































































Chapter 10 

DESCRIPTION OF ERODED GUN BORES 

By Lloyd E. Line , /r. a 


101 DEFINITION OF TERMS 

E nlargement of the bore of a gun by erosion rad¬ 
ically affects the flight of the projectile. In gen¬ 
eral, erosion at the breech end decreases the range of 
the gun and erosion at both ends decreases its accu¬ 
racy. For these reasons it is important to seek means 
of mitigating or controlling bore enlargement. 

This chapter contains a description of eroded guns 
with respect to the superficial effects of the eroding 
process that are deduced from physical measure¬ 
ments of bore dimensions and from visual examina¬ 
tion of the surface. More fundamental effects of gun 
erosion will be described in Chapter 12. 

For the purposes of this Summary Technical Re¬ 
port, the term “gun erosion” is defined as the gradual 
changes in bore dimensions as a result of normal fir- 
ing b and the changes in character of the bore surface 
and walls that lead to these changes in dimensions. 
Such a definition would satisfy the metallographer 
who might detect a slight flow of gun metal or the 
presence of a thermally altered layer after the firing 
of only one round. It would also satisfy a star gauger 
who measures changes in bore diameter or a ballisti- 
cian who notes the effect of such changes on gun per¬ 
formance. 

In any case, by “gun erosion” we mean a result and 
not the process that produces the result. In Sections 
10.3 and 10.4, entitled “Origin Erosion” and “Muzzle 
Erosion,” respectively, the term “erosion” is used in 
a macroscopic sense; that is, it is synonomous with 
the enlargement observable with the eye or Avith star 
and plug gauges. In Section 10.5, entitled “Nature 
of the Eroded Surface,” many of the microscopic 
changes in the bore surface, such as liquefaction of 
surface material and pebbling, will also be included 
in the term “erosion” because these changes definitely 
lead to the macroscopic changes in dimensions that 
ultimately affect gun performance. 


a Technical Aide, Division 1, NDRC. (Present address: 
Chemistry Department, University of Tennessee, Knoxville, 
Tenn.) 

b This definition is not intended to include the increase in 
bore diameter that results from excess powder pressure, either 
during proof firing or later. 


If we measure the bore diameter near the origin of 
rifling across several of the lands of a gun that has 
been fired an appreciable number of rounds, we find 
that this diameter is greater than that of a new gun; 
a similar result will be noted for the grooves. We ex¬ 
press this land erosion and groove erosion as the diam¬ 
etral enlargement across the lands and grooves, 
respectively. 

The erosion of a gun is not uniform along its length 
but is localized in the region of the origin of rifling in 
all guns, and it is localized also in the region of the 
muzzle in many guns (particularly those of relatively 
high velocity). The former type of erosion we call 
origin erosion, the latter we call muzzle erosion. As is 
described in detail in Sections 10.3.2 and 10.4.3, the 
origin erosion decreases toward the muzzle and the 
muzzle erosion in most cases increases toward the 
muzzle, so that in practically all cases origin erosion 
and muzzle erosion occur in two distinct regions. Be¬ 
tween them there frequently occurs a region several 
calibers long where the bore is constricted, which is 
usually attributed to coppering, that is, the deposi¬ 
tion of material from the rotating band of the pro¬ 
jectile. This phenomenon is described briefly in Sec¬ 
tion 10.5.4. 

In origin erosion we find both land erosion and 
groove erosion, the former being some two or three 
times as great as the latter; in muzzle erosion we find, 
for all practical purposes, only land erosion. Figure 1 
shows how the land and groove erosion vary along the 
bore for the 4.7-in. gun, T2, No. 2. The origin erosion 
here is typical of guns in general. Not only do we 
speak of erosion as general enlargement of a region of 
the bore, but we often wish to speak of erosion at a 
particular position along the bore. Thus we speak of 
erosion at the origin of rifling, erosion at the muzzle, and 
so forth. 

As will be seen later (Sections 10.3.6 and 10.4.8), 
erosion at a particular position is not always circu¬ 
larly symmetrical. In referring to this fact we employ 
the term asymmetric erosion. Muzzle erosion is usu¬ 
ally asymmetric. (In the literature, the terms oval, 
elliptic, and eccentric are often used to indicate asym¬ 
metry. Star gaugers at Army establishments use the 
term out-of-round?) 


CONFIDENTIAL 


195 



INCREASE IN DIAMETER ACROSS GROOVES IN INCHES INCREASE IN DIAMETER ACROSS LANDS IN INCHES 


196 


DESCRIPTION OF ERODED GUN BORES 




Figure 1. Star-gauge curves for 4.7-in. gun. 


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MEASUREMENT OF GUN EROSION 


197 


Asymmetric erosion gives configurations, such as 
an oval, a circle with a localized region of erosion 
(pocket), or perhaps a circle eccentric with the orig¬ 
inal bore. In any case the configuration has some 
orientation (such as 7 o’clock, breech time) with re¬ 
spect to the mounting of the barrel. If, along the bore, 
the orientation of the configuration twists more or less 
with the rifling, as seems to be the case for muzzle 
erosion, we say the erosion “follows the rifling” or is 
of the spiral type. In some large guns spiral erosion is 
immediately evident to the observer. 

The high cost of guns and of ammunition has pre¬ 
cluded the firing of any large number of medium or 
large-caliber guns just for the sake of obtaining ero¬ 
sion data. Even in the case of small arms extensive 
testing under a large variety of conditions was not 
carried out until World War II. Then the diversion of 
an insignificantly small proportion of the millions of 
caliber .50 machine gun barrels being produced pro¬ 
vided ample quantities for testing purposes. Such tests 
were performed at Purdue University for the Army 
Ordnance Department 294 as well as at Aberdeen Prov¬ 
ing Ground. In Great Britain a long series of firings 
under a great many different conditions was made 
under the auspices of the Barrel Life Panel. 405 406 

10 2 MEASUREMENT OF GUN EROSION 
10 21 Star Gauges 

The ways in which a change in bore diameter may 
be brought about are (a) removal or displacement of 
bore surface material, (b) distortion of idfling, (c) bar¬ 
rel expansion or contraction, and (d) coppering. A 
star gauge gives no indication of the predominant 
change; it merely registers a change in diameter due 
to one or more of the above causes. It is therefore well 
to bear in mind these considerations in studying star 
gauge data. As will be seen later, measurement of the 
bore diameter with plug gauges also has certain limi¬ 
tations. 

At Army and Navy establishments gun erosion and 
bore enlargement in general (for cannon) are measured 
by a star gauge. This instrument measures the bore di¬ 
ameter of a gun across lands and across grooves. The 
points of the gauge are extended radially in a plane 
by the movement of a handle at the opposite end of 
a long staff until they contact the surfaces of the 
lands (or the grooves, as the case may be). The di¬ 
ameter is read from a vernier scale on the handle. 

There are several kinds of star gauges, the details 


of which may be found elsewhere. 288 Essentially they 
may be classified according to whether they have two 
measuring points 180 degrees apart or three measur¬ 
ing points 120 degrees apart. At Army establishments 
these are called two-point and three-point gauges, re¬ 
spectively. 

It is the practice at Army establishments to make 
two measurements of diameter with a star gauge in 
order to take asymmetric erosion into account. In the 
case of a two-point star-gauge, the diameter is mea¬ 
sured in two planes perpendicular to each other, and 
the two measurements are averaged. This average 
represents contact by the points with the lands (or 
grooves, as the case may be) at four places, 90 degrees 
apart. Similarly, with the three-point star gauge, the 
diameter is measured, first with the “Y” in one posi¬ 
tion, and then with the “Y” rotated 180 degrees, and 
the two measurements are averaged. This average 
represents contact by the points of this gauge with 
the lands (or grooves) at six places, 60 degrees apart. 
Thus the measurement of erosion by a star gauge 
yields a fairly average diametral change at a particu¬ 
lar position along the bore. 

At Army establishments, bore diameters are usu¬ 
ally measured at a number of positions with intervals 
no greater than 5 in. It is the practice to star gauge 
the gun from muzzle to the breech with the measur¬ 
ing points of the star gauge resting on the same lands 
or grooves, as the gauge moves down the bore. A plot 
of the diameter or increase of diameter of a gun as a 
function of distance from some fixed point, such as 
the origin of rifling, is called a star gauge curve. At the 
Naval Proving Ground such a curve for the lands is 
called a bore profile. Usually a star gauge curve repre¬ 
sents the average diametral increase given by the 
gauge in two positions, as previously described. 
Sometimes, however, the “horizontal” and “vertical” 
or the “Y-up” and “Y-down” readings are plotted 
separately to show asymmetric erosion. These posi¬ 
tions, at Army establishments, have reference to the 
position of the gauge when a reading is taken at the 
origin of rifling. 

Most star gauges can be read accurately to 0.001 
in., and an experienced operator can duplicate his 
measurements with an error no greater than 0.001 in. 
But among different operators, including experienced 
and inexperienced ones, there might be variations of 
0.005 in. or more in the measurement of the same 
diameter. 

A limitation of the ordinary star gauges, especially 
with regard to muzzle erosion, should now be pointed 


CONFIDENTIAL 




198 


DESCRIPTION OF ERODED GUN BORES 


out. As previously mentioned, erosion is often asym¬ 
metrical. Since a star gauge measures a diameter in¬ 
stead of a radius, it gives no information concerning 
erosion at different positions around the periphery. 
Furthermore, the average of the diameters indicated 
by the two positions of the star gauge may be some¬ 
what different from the true average diameter that 
would be given by an average of the readings with the 
gauge in a large number of positions around the bore. 
Especially may this be true in cases where the asym¬ 
metry is characterized by erosion in a rather narrow 
region of the periphery. 


10 2 2 Plug Gauges 

Another way in which erosion is measured is by 
means of the advance of a plug gauge having a cylin¬ 
drical, conical, or other type of head. Such a gauge 
has been used both to obtain the bore profile of a gun 
and to indicate the end of life. 

In studies with the Franklin Institute caliber .50 
erosion-testing gun (Section 11.2.1) plug gauges hav¬ 
ing cylindrical heads ranging from 0.490 in. to 0.516 
in. in diameter (in steps of 0.002 in.) were used to 
measure diameters across lands. Also rifled plug 
gauges ranging from 0.511 in. to 0.529 in. in diameter 
(in steps of 0.002 in.) were used to measure diameters 
across grooves. Thus gauge diameters plotted against 
distance of the head from the breech gave the profile 
of the bore. 

Such gauges were found to have the advantages of 
simplicity and speed, but they could be used only in 
measuring the erosion of materials, such as gun steel, 
the erosion of which produced a bore that tapered in 
only one direction. They were not reliable in measur¬ 
ing the erosion of bore surfaces which had been plated 
with a resistant coating, such as molybdenum or 
chromium (Section 16.4). A small block of resistant 
material adhering to the surface (beyond which meas¬ 
urable erosion has occurred) easily stops the advance 
of a plug gauge. 122 

Because of erosion in the neighborhood of the ori¬ 
gin of rifling, the forcing cone does not retain its orig¬ 
inal position; it gradually moves forward. The ad¬ 
vance of the forcing cone is measured with a special 
tapered plug gauge that is inserted through the 
breech until it is stopped by contact with the “new” 
forcing cone. 288 Accompanying the advance in forcing 
cone is a loss of muzzle velocity, so that such an ad¬ 
vance gives some indication of the gun performance 


with respect to muzzle velocity. (See Section 10.3.7 
and 23.1.4.) 

As the measurement of the forcing cone advance is 
much simpler than a star-gauge measurement, it is 
more readily performed in the field. Although the re¬ 
lation between muzzle velocity loss and forcing cone 
advance is nonlinear for most guns, it nevertheless 
can be used as a criterion of the end of life. This was 
done in World War I by the French and subsequently 
by the AEF when using French guns. 16 During World 
War II studies were carried out by the Army and 
Navy on cannon with gauges having different types 
of heads. 0 


10 3 ORIGIN EROSION 

10 31 Relative Erosion of Lands and Grooves 

In general, land erosion is greater than groove ero¬ 
sion at any stage in gun life or at any position along 
the bore. Usually it is two or three times as great in 
medium and large caliber guns. The relative rates of 
erosion, however, seem to depend on land height and 
on degree of land erosion. At the origin of rifling the 
lands at first erode much faster than the grooves, but 
the rates tend to equalize and become nearly uniform 
as the land and groove diameters approach equality. 

In seven barrels of five different calibers, 37-mm to 
8-in., selected because of best uniformity of firing, the 
lands, which varied in height radially from 0.02 to 
0.14 in., eroded down to their original bases while the 
grooves deepened almost a fixed amount (0.020 ± 
0.005 in., radially). The subsequent rates of erosion 
of the lands of each gun were scarcely greater than of 
the grooves and both were nearly uniform. Possibly 
in 37-mm guns and in ones of smaller caliber, under 
“mild” conditions of firing, the lands and grooves 
would erode almost equally from near the start. In 
guns of larger caliber the three-sided exposure of the 
high-standing lands above relatively narrow grooves 
would be expected to favor rapid erosion. 75 

The prime importance of the protrusion of the 
lands into the gas stream on heat transfer, as related 
to flame temperature and thermally altered layers, 
which was shown in studies 124 with the caliber .50 
erosion-testing gun (Section 11.2.1), is discussed in 
Section 15.3.3. 


c Information received by the author on visits to Aberdeen 
Proving Ground and the Naval Proving Ground. 


CONFIDENTIAL 




ORIGIN EROSION 


199 


10 3 2 Variation Along the Bore 

In cannon, the erosion of both lands and grooves is 
highest at the origin of rifling. Toward the muzzle it 
drops off rapidly at first, with the rate of change be¬ 
coming progressively less, so that at several calibers 
from the origin of rifling there is no measurable en¬ 
largement, but cracking is extensive. In most cases 
this variation along the bore is regular; that is, a 
smooth curve can be drawn through the points repre¬ 
senting increases in diameter, and the curves of all 
guns are quite similar in form. Typical star-gauge 
curves for origin erosion are given in Figure 1. In 
chromium-plated guns the extension of the origin ero¬ 
sion toward the muzzle is less than in nonplated ones, 
other things being equal. Ordinarily, groove erosion 
does not extend as far as land erosion. 

For equivalent rounds fired in the caliber .50 
erosion-testing gun (Section 11.2.1), the extension of 
land erosion with hot (double-base) powders was 
greater than with cooler (single-base) powders. 123 The 
extension of land erosion forward from the origin of 
rifling in chromium-plated guns is less than it is for a 
nonplated gun after the same number of rounds, as 
evidenced by the fact that after 83 rounds with 
double-base powder, the extension in 90-mm guns for 
chromium-plated bores was about 10 in. from the ori¬ 
gin of rifling whereas it was about 50 in. for the non¬ 
plated ones. There was little difference in the magni¬ 
tude of erosion at the origin of rifling between the 
plated and nonplated guns. d 

There is often conspicuous erosion in guns behind 
the origin of rifling. This has received but little atten¬ 
tion. Because the thermal factors in the cause of ero¬ 
sion are more important here than the mechanical, 
as is brought out in Section 13.2.5, chromium plate 
offers considerable protection. 

In a case gun, erosion between the position of the 
mouth of the case and the origin of rifling is often 
about as much as in the adjacent grooves. In a badly 
worn bag gun, there is a gradual decrease of erosion 
rearward one to three calibers from the origin of rifling. 
An eroded, pebbled surface gives place to a cracked 
surface. 16 

For small arms fired in bursts, origin erosion ex¬ 
tends over a much greater proportion of the barrel 
than in cannon, particularly if the bursts are long. 
For example erosion extended to the muzzle in a cal- 


d Author’s observation from data supplied by Aberdeen 
Proving Ground. 


iber .50 machine gun after it had been fired one con¬ 
tinuous burst of 250 rounds. 406 Serious effects result¬ 
ing from the weakening, and expansion and contrac¬ 
tion, of the whole barrel wall as a result of continued 
fire are the subject of Sections 5.6.4 and 10.5.3. 

10 3 3 Development of Origin Erosion 
with Firing 

The extent of origin erosion in a gun is ordinarily 
indicated numerically by the increase in diameter of 
the lands Ad 0 at the origin of rifling or immediately 
in front of it. In standard nonplated guns of medium 
or large caliber firing banded projectiles the land 
erosion at the origin of rifling and positions forward 
of it begins perhaps with the first round, its rate being 
higher initially than during later rounds. For 14-in./ 
50-cal. guns this variation of erosion A d 0 with the 
number of rounds N has been expressed by the Bureau 
of Ordnance, Navy Department, as an exponential 
function 16 according to equation (1). 

Ado = 0.349 (1 - e -o.oio65*)_ (i) 

Studies of such curves have shown that in some 
guns the erosion rate eventually becomes nearly con- 



Figure 2. Erosion at origin of rifling versus number of 
rounds for 4.7-in. gun. 


stant, 75 as exemplified by Figure 2, for 4.7-in. gun, 
T2, No. 2. 

A cursory examination by the author of curves of 
many Naval guns in which origin erosion was plotted 
against E.S.R. (equivalent service rounds) e showed a 

e In the U. S. Navy for the purpose of tabulation, warming 
rounds or any other rounds fired at other than the standard 
charge are converted into equivalent service rounds by means 
of the formula E.S.R. = ( W/W s ) 6 , where W is the weight of 
the charge actually used and W s is the weight of the service 
charge. 


CONFIDENTIAL 






200 


DESCRIPTION OF ERODED GUN BORES 


falling off of the erosion rate with E.S.R. and straight 
portions of curves beginning at a value of land ero¬ 
sion roughly twice the land height. The attainment 
of a constant rate is considered in Section 10.3.1. 

That the erosion rate eventually becomes constant 
was shown also by studies of erosion with the caliber 
.50 erosion-testing gun. 122 The progress of erosion ob¬ 
served in some of these tests is given in Section 15.3.1. 
A study of erosion rates using pre-engraved projec¬ 
tiles (Chapter 31) showed that once erosion started, 
the rate was about the same as that shown by the 
straight portion of the erosion curve for banded bul¬ 
lets of the artillery type. This result led to a rough 
distinction between gas erosion and erosion caused by 
the forces of engraving, e.g. friction and swaging (see 
Section 13.4.2). Thus the feature of the gradual di¬ 
minishing of the erosion rate of the lands to a con- 


O 

z 



Figure 3. Extension of origin erosion for 4.7-in. gun. 

stant value during the life of a gun may be related to 
a diminishing of the mechanical forces in engraving, 
as the amount of the engraving lessens in the worn 
gun. 

Not only does the origin erosion rate decrease with 
the number of rounds, but also the same holds true 
roughly for the rate of forward extension of the ero¬ 
sion. A typical example of this variation is given in 
Figure 3 for 4.7-in. gun tube, T2, No. 2. 

Thus we see that in origin erosion a sort of “cone” 
of wear develops, extending toward the muzzle from 
the origin of rifling. Both the “base” and the “height” 
enlarge with firing, at first rapidly and later more 
slowly. In many guns, while origin erosion is develop¬ 
ing, muzzle erosion is developing in the forward por¬ 
tion of the barrel. (This will be described in Section 
10.4.) In connection with the falling off of the rate of 
general erosion there is a similar occurrence with re¬ 
spect to velocity loss. Thus, Figure 4, which is typical 
of guns in general, shows that the velocity drops at 
first rapidly and later more slowly. 


10 3 4 Dependence on Caliber 

The erosion per round and also the maximum of 
erosion at the end of life increase very greatly with 
increase of caliber. Representative values of the for¬ 
mer quantity for several guns of about the same muz¬ 
zle velocity but different calibers are given in Table 
1. It should be remembered that the variation in the 


Table 1. Comparison of average erosion per round for 
guns of different caliber: velocity, 2,500 to 2,750 fps. 16 





Erosion Rate* 

Gun 

Rounds 

Lands 

Grooves 

37-mm M3 Tube No. 22708 

902 to 1472 

0.000047 

0.000026 

3-in. AA Gun M3 No. 640 

326 to 

528 

0.00017 

0.00012 

8-in. M1888 Mil No. 56 

62 to 

127 

0.00035 

0.00027 

14-in. Ml920 Mil No. 11 

123 to 

154 

0.0011 

0.00096 

16-in. M1919 Mil No. 2 

123 to 

152 

0.0028 

0.0014 


♦Average erosion rate at origin of rifling in inches per round. 


rate of erosion of guns of the same caliber and muzzle 
velocity is quite large, and that therefore the values 
given in Table 1 should not be considered as estab¬ 
lishing functional relation between erosion and caliber. 
They are merely illustrative of possible values. 

A functional relationship between erosion or erosion 
rate and caliber might possibly be established if data 
on guns (fired in a uniform manner) that are more or 
less scale models of each other could be assembled. 



ROUNDS 

Figure 4. Muzzle velocity versus number of rounds 
for 90-mm guns. (Average of several guns.) 

10 3-5 Dependence on Muzzle Velocity 

It has long been known that the rate of origin 
erosion increases very rapidly with an increase in 
muzzle velocity. This fact has prohibited the employ¬ 
ment of muzzle velocities for standard guns much be¬ 
yond about 3,000 fps. Some idea of the relation of 
muzzle velocity to origin erosion may be obtained 
from Table 2. 


CONFIDENTIAL 











ORIGIN EROSION 


201 


Table 2. Relation of muzzle velocity to origin erosion.* 


Proj. Service Average origin erosion f 

wt. Charge velocity Nonplated guns Chromium plated guns 


Gun 

(lb) 

(lb) 

(fps) 

400 E.S.R. 

800 E.S.R. 

400 E.S.R. 

800 E.S.R. 

5-in./25-cal. 

53.55 

9.75 

2175 

0.055(1) 

0.090(1) 

0.041(5) 

0.068(5) 

5-in./38-cal. 

54.9 

15.0 

2600 



0.076(94) 

0.116(94) 

5-in./51-cal. 

50.0 

27.0 

3150 

0.141(7) 

0.225(1) 

0.127(12) 

0.155(6) 


* These data were furnished by the Naval Proving Ground. 

t A number in parenthesis indicates the number of guns for which the value was averaged 


The erosion rate, of course, is not fundamentally 
related to muzzle velocity. Most likely, it is related to 
the powder charge, which ordinarily is increased to 
increase the velocity. This may be seen from the fact 
that there was little difference in the erosion of a 
45-in. barrel (muzzle velocity 3,458 fps) and of an 85- 
in. barrel (muzzle velocity 4,200 fps) when fired with 
about the same charge (400-405 grains) of double¬ 
base (FNH-M2) powder and pre-engraved projec¬ 
tiles in the caliber .50 erosion-testing gun. The rate 
of land erosion at 0.5 in. from the origin of rifling was 
of the order of 20 X 10 -5 in. per round. 122 

Aside from the size of the powder charge, erosion 
rate depends also on the type of powder used, as evi¬ 
denced by the fact that different propellants fired in 
the caliber .50 erosion-testing gun at the same muzzle 
velocity (3,300 ± 50 fps) gave different erosion 
rates. 123 This subject is dealt with in Chapter 15. 

10 3 6 Asymmetry of Erosion 

Origin erosion is usually not circularly symmetri¬ 
cal, although asymmetry of origin erosion is ordinar¬ 
ily less than that of muzzle erosion. It is usually de¬ 
tected in the course of star gauging by the difference 
between “horizontal” and “vertical” or “Y-up” and 
“Y-down” readings. Table 3 is illustrative of asym¬ 
metry encountered in Service guns. The writer has 
observed that in the majority of cases, especially in 


Army 155-mm guns, Ml, the vertical erosion is 
greater than the horizontal erosion. 


10 3 7 Effect of Origin Erosion on 
Gun Performance 

As has been mentioned, origin erosion decreases 
the range and accuracy of a gun. To some extent 
these effects are interrelated, but they will be separ¬ 
ated for purposes of discussion. 

The decrease in range is due in a large measure to 
the loss in muzzle velocity that accompanies origin 
erosion. Such a loss is progressive, its rate being 
greater in the early stages of firing than in the later 
stages. A typical curve showing the variation of the 
muzzle velocity with the number of rounds for several 
90-mm guns is given in Figure 4. 

Perhaps the most important effect of origin erosion 
on muzzle velocity is that the projectile is more eas¬ 
ily moved forward during the early stages of burning; 
that is, the starting pressure [Po in equations (3a) 
and (3b) of Chapter 3] is less than for a new bore. As 
a result, the rate of burning of the powder is changed 
in such a way that the average pressure developed 
over the length of the bore is less, with a consequent 
loss in muzzle velocity. Moreover, in a bag gun the 
projectile is rammed farther forward than normally, 
thus reducing the maximum pressure. 


Table 3. Asymmetry of origin erosion in several Service guns. 


Erosion (in.) 

Rounds Lands Grooves 


Gun 

Designation 

fired 

“Horizontal” 

“Vertical” 

“Horizontal” 

“Vertical” 

4.7-in. 

T2, No. 3 

93 

0.061 

0.057 

0.019 

0.027 

155-mm 

M1A1, No. 18 

1148 

0.057 

0.086 

0.016 

0.032 

6-in. 

B.L. Mk XXIII* 


0.100 

0.075 



8-in. 

Ml, No.20 

572 

0.250 

0.263 



12-in. /45-cal. 

Mk V 

183 

0.350 

0.386 




♦British gun. 


CONFIDENTIAL 











202 


DESCRIPTION OF ERODED GUN BORES 


It has been suggested 509 that leakage of gas past 
the projectile accounts in some measure for pressure 
and velocity loss, but this has been questioned. 515 
However, tests in the caliber .50 erosion-testing gun 
with obturated and unobturated pre-engraved pro¬ 
jectiles and projectiles of the base-cup type, which 
are described in Section 31.4.2, seem to indicate that 
gas leakage is important in this respect. 122 

Loss in range is also due to interior ballistic condi¬ 
tions caused by origin erosion. For example, the or¬ 
igin erosion may reach a stage wherein the lands are 
removed over a considerable length of the barrel. As 
a result the rifling may fail to impart sufficient rota¬ 
tional spin to the projectile causing it to wobble in 
its flight with consequent loss in range. Also, the ro¬ 
tating band may be sheared instead of engraved, so 
that practically no spin is imparted to the projectile. 

As a gun is fired, its muzzle velocity diminishes, so 
that its accuracy diminishes progressively in the 
sense that it cannot hit the target at a given distance. 
However, in addition, the erosion apparently gives rise 
also to a dispersion of the shots. Usually the disper¬ 
sion is greater in range than in azimuth. 

Part of the dispersion in range is related to the dis¬ 
persion in the muzzle velocity due to erratic burning 
of the powder and other interior ballistic effects. In 
addition, the geometry of the worn gun has an effect. 
As previously mentioned, the rifling may be so worn 
that insufficient spin or no spin at all (due to strip¬ 
ping of the rotating band) may result for a large num¬ 
ber of shots. Also, a worn bore may allow gas to leak 
by the projectile in an indeterminate fashion from 
round to round causing the bullet to tip in the bore 
with consequent dispersion. 1 In the caliber .50 erosion- 
testing gun (Section 11.2.1) the base-cup type of bul¬ 
let (designed for better obturation) gave a mean radi¬ 
us of dispersion at 45 ft of 0.6 in. whereas the stan¬ 
dard ball bullet, M2 gave a dispersion of 2.2 in. under 
the same test conditions. 122 Furthermore, the worn 
condition of the bore, especially if it extends forward 
appreciably, may allow the projectile to tip or ballot 
in the bore so that it emerges from the muzzle with 
an initial yaw. 

While velocity loss occurs progressively, excessive 
dispersion develops more or less abruptly. This is es¬ 
pecially true in the burst firing of caliber .50 machine 
guns in which keyholing occurs rather suddenly. Part 
of this, however, is attributable to the expansion of 


f Gas leakage is also believed to be responsible for some cases 
of body engraving. (See Section 10.4.10.) 


the gun due to the heating caused by the burst firing. 
Because of this expansion, the effect of temperature 
on gun performance, discussed in Section 5.6.3, is 
similar to the effect of erosion. The sequence of events 
as the machine gun barrel wall weakens during rapid 
fire is given in Section 5.6.4. 

At the Naval Proving Ground range dispersion of 
a gun is said to develop rather abruptly in the course 
of firing. 102 

10.3.8 Erosion Ahead of a Liner Joint 

The use of short breech gun liners (discussed in 
Parts VI and VII) has added a new feature to gun 
erosion. When a gun containing a steel liner at the 
breech is fired, the bore diameter increases just ahead 
of the liner as well as at the origin of rifling. This 
phenomenon, which is attributed to increased gas 
turbulence, was studied first with the caliber .50 
erosion-testing gun (Section 11.2.1). 

Two liners were inserted in tandem in the barrel so 
that two liner joints at different places in the bore 
would be exposed simultaneously to the powder gases 
from the same charge. The main purpose was to de¬ 
termine the effectiveness of different methods of seal¬ 
ing the joint so that the powder gases would not pen¬ 
etrate behind the liner. A lapped joint was found to 
be preferable to one employing a gasket. In addition 
it was observed that erosion was much greater just 
beyond the liner joint than just behind it. 

The same effect was observed during the firing of 
the 90-mm gun, T19, containing a replaceable steel 
liner (Section 26.3). As may be seen from Figure 8 of 
Chapter 26, erosion just beyond the joint was appre¬ 
ciable long before the origin erosion had extended as 
far forward as the liner joint. The rate of this erosion 
was a linear function of the total number of rounds 
fired, independent of the condition of the origin of 
rifling. This result clearly demonstrated that this 
type of erosion is almost entirely attributable to the 
powder gases. 

A machine gun barrel containing an erosion- 
resistant liner presents a different situation with re¬ 
spect to erosion beyond the liner joint. In this case 
the liner makes it possible to fire the gun so much 
longer than is possible with a plain steel barrel that 
powder-gas erosion of the steel section beyond the 
liner presents a serious problem. The enlargement of 
the bore and obliteration of the rifling there contri¬ 
bute td the eventual failure of the barrel through 
both inaccuracy and loss of muzzle velocity. After the 


CONFIDENTIAL 




MUZZLE EROSION 


203 


enlargement becomes great enough for blow-by of the 
powder gases (Section 13.4.1), the increased heating 
of the bore surface further accelerates the erosion 
there and also increases the undesirable effects on 
ballistics from general heating of the barrel wall, 
discussed in Section 5.6.4. 


104 MUZZLE EROSION 

1041 Introduction 

The importance of muzzle erosion has become ap¬ 
parent only in recent years. While origin erosion 
occurs in all guns, muzzle erosion occurs only in some 
guns, particularly those of relatively high muzzle ve¬ 
locity. This kind of erosion has not been studied as 
thoroughly as origin erosion. One reason for this is 
that in most instances it is not nearly so conspicuous 
as is origin erosion; hence its effects on the life of pres¬ 
ent Service weapons have been considered of second¬ 
ary importance, if any. However, it seemed to Divi¬ 
sion 1, whose ultimate goal was the achievement of 
hypervelocity guns, that muzzle erosion might well 
limit the life of a gun firing at a muzzle velocity in 
excess of 4,000 fps even though the problem of origin 
erosion were solved. 

In order to have a body of knowledge that would 
clarify this situation and perhaps point to possible 
ways of eliminating muzzle erosion, a survey 67 of 
existing information was made. Following this gen¬ 
eral survey, a collection 102 of previously unpublished 
data at the Naval Proving Ground was made. Then 
a careful study was made of the bore surface of the 
muzzle sections of some typically eroded guns: a 
57-mm Army gun, an 8-in. Army gun, and a 14-in. 
Naval gun liner. The results 130 supported the conclu¬ 
sion, tentatively stated in the earlier survey, 67 that 
the mechanism of muzzle erosion is different from 
that of origin erosion in that, whereas the primary 
mechanism of the latter involves the powder gases, 
that of the former involves mechanical action of the 
projectile (chiefly rubbing) on the muzzle lands. 

The following description of muzzle erosion, which 
parallels that of origin erosion in Section 10.3, is 
based on the three reports 67 ’ 102 ’ 130 just mentioned. 
The very paucity of the previous information on this 
subject makes it desirable to present the available 
data here in more detail than for origin erosion, for 
the benefit of those who are likely to undertake stud¬ 
ies of muzzle erosion in the future. 


10,4 2 Relative Erosion of Lands and Grooves 

The relative erosion of lands and grooves at the 
muzzle differs greatly from that at the origin of ri¬ 
fling. At the muzzle, when the increase in diameter 
across the lands is small, say of the order of 10 -3 in., 
the corresponding increase in diameter across the 
grooves is of the same order of magnitude. But when 
the increase in diameter across the lands is relatively 
large, say of the order of 10“ 2 or 10 _1 in., the corre¬ 
sponding increase across the grooves is still no greater 
than 10“ 3 in. It is sometimes negative, owing pre¬ 
sumably to copper from the rotating band. Thus, 
groove erosion at the muzzle is practically negligible. 
Therefore muzzle erosion ordinarily refers to erosion 
of the lands. These tendencies are shown by Tables 
4 and 5. 

Table 4. Land erosion at the muzzle versus correspond¬ 
ing groove erosion when land erosion is slight, as indi¬ 
cated by measurements of average bore diameter with a 
star gauge. 67 Measurements taken at or very near the 
muzzle face. 


Num- Erosion at the 

U un _ berof muzzle, change 

Desig- Num- rounds in diameter (in.) 


Caliber 

nation 

ber 

fired 

Lands 

Grooves 

37-mm 

M1A2 

313 

5375 

0.005 

0.003 

37-mm 

M3 

2110 

6082 

0.004 

0.004 

90-mm 

Ml 

35 

302 

0.006 

0.005 

4.7-in. 

T2E1 

2 

188 

0.004 

0.004 

155-mm 

M1918M1 

842 

812 

0.007 

0.005 

12-in. 

M1895 

40 

157 

0.006 

0.004 


Table 5. Land erosion at the muzzle versus correspond¬ 
ing groove erosion when land erosion is large as indi¬ 
cated by measurements of average bore diameter with a 
star gauge. 67 Measurements taken at or very near the 
muzzle face. 


Num- Erosion at 



Gun 


ber of 

muzzle, change 

Caliber 

Desig¬ 

nation 

Num¬ 

ber 

rounds 

fired 

in diameter (in.) 
Lands Grooves 

57-mm 

Ml 

1941 

556 

0.042 

-0.001 

3-in. 

M3 

909 

1269 

0.017 

0.001 

6-in. 

Brown 


88 

0.045 

-0.012 

8-in. 

Mk VI, 

Mod 3A2 

173L2 

295 

0.054 

-0.002 


10 4 3 Variation Along the Bore 

Muzzle erosion is usually found over a considerable 
length of the bore at the muzzle end of the gun— 
some 15 or 20 calibers in most cases. Unlike origin 


CONFIDENTIAL 














204 


DESCRIPTION OF ERODED GUN BORES 


erosion, its variation along the bore is not uniform 
among different guns. The shapes of the curves are 
varied. Some of the types that have been observed 
are shown in Figure 5. Curve D illustrates the ex¬ 
treme case of measurable erosion throughout the 
length of the bore. 








DISTANCE FROM ORIGIN OF RIFLING 


Figure 5. Types of bore profiles. (Schematic.) 


In terms of diametral change, muzzle erosion is al¬ 
most never as great as origin erosion. But in terms of 
total volume change, as judged from the areas under 
the star gauge curves, the erosion at the muzzle end 
is often greater than that at the breech end. In most 
worn guns, the muzzle erosion decreases less rapidly 
toward the breech than does origin erosion toward the 
muzzle. 


10 4 4 Development of Muzzle Erosion 
with Firing 

For most guns, the maximum muzzle erosion is at 
or very near the muzzle face. This quantity is gener¬ 
ally taken as a measure of the degree of muzzle ero¬ 
sion. In this chapter it is denoted by A d m where d is 
the bore diameter and m refers to the muzzle. 


Information received from Naval Ordnance per¬ 
sonnel based on erosion data accumulated at the 
Naval Proving Ground indicates that A d m is approx¬ 
imately a linear function of the number of actual 
rounds N, beginning with the first round, provided 
the charge for each round is the same. Thus about the 
same amount of muzzle erosion may be said to ac¬ 
company the firing of each round, and we may as¬ 
cribe to the muzzle erosion of a particular gun a char¬ 
acteristic rate A d m /N, which is more or less indepen¬ 
dent of the stage of life of the gun. Some exceptions 
to this generalization have been found. 

The linear relationship is said to apply also to 
chromium plated guns, but in these guns the muzzle 
erosion is delayed by the presence of chromium so 
that the linear function does not begin with the first 








Figure 6. Variation of erosion at the muzzle versus 
E.S.R. for Naval guns. 


round. It is said, however, that when the muzzle ero¬ 
sion begins, its rate is very nearly the same as that of 
the corresponding nonplated gun. (See Figure 6A.) 

Plots of muzzle erosion vs E.S.R., when the charges 
are not the same for each round, do not give straight 
lines through the origin of coordinates. The curves 


CONFIDENTIAL 
























MUZZLE EROSION 


205 


are varied in form. This may be due to coppering of 
the muzzle end when low-charge rounds are fired. 

Thus, it has been observed at the Naval Proving 
Ground that in guns fired with mixed rounds, i.e., 
groups of service rounds alternated with groups of 
low-charge rounds, the muzzle erosion is slight. If 
after firing mixed rounds, service rounds only are 
fired, the muzzle erosion becomes appreciable, its rate 
being the same as would be the case if the gun had 
fired only service rounds. If the firing of mixed rounds 
is resumed, the muzzle enlargement actually de¬ 




creases. These phenomena, illustrated schematically 
by Figure 6B, are explained by the assumption that 
the firing of low-charge rounds tends to copper the 
muzzle, which protects it from erosion, and the firing 
of service rounds removes the copper, thus exposing 
the surface to erosive influences. 102 

Studies of star gauge curves have shown that muz¬ 
zle erosion is not usually confined to a relatively short 
section of one or two calibers, as is often thought to 
be the case. In fact, it may extend as far back as 15 to 
30 calibers from the muzzle. Furthermore, such ex- 





/ 

x-" 


/ 

/ 

/ 



f 

1 

- 1 

1 



1 

J 

1 

J 



j 14-IN./50 CAL. 
t NAVY GUN MKC 

NO. 110 L 2 

f 

I_i_ 

_ 1 _ 

_ 1 _ 


- 

• 


o 

\ 

so o 



OO o < 


-7 

/ < 

- / 

/ 

i 

) 



/ 8-IN.ARMY GUNS MICE! 

/ MOD 3A2,T2 AND T2E1 

i 

l 

_i_ 



_ l _ 


0 0.01 0.02 0.03 0.04 


Ad m IN INCHES 

Figure 7. Correlation of extension toward the breech of muzzle erosion with A d m . (This figure has appeared as 
Figure 7 of NDRC Report No. A-357.) 


CONFIDENTIAL 




















































206 


DESCRIPTION OF ERODED GUN BORES 


tension of erosion seems to be reached relatively 
quickly, as exemplified by Figure 7. Thus when A d m 
for the 8-in. guns was only 0.01 in., erosion could be 
detected as far back as 15 calibers from the muzzle. 
But with further increase in diameter up to 0.03 in. 
there was little extension of the erosion. 

10 4 5 Dependence on Caliber 

It is usually said that muzzle erosion occurs only in 
large-caliber guns. While it is true that it is most con¬ 
spicuous in those guns, examination of star gauge 
data shows that some muzzle erosion occurs in prac¬ 
tically all guns. It has been shown 67 that the muzzle 
erosion rate Adm/N varies approximately as the 
square of the bore diameter for standard guns of dif- 


Ad m (IN) INCREASE IN AVERAGE BORE DIAMETER 
ACROSS LANDS AT MUZZLE 
N NUMBER OF ROUNDS FIRED 
d(IN) NORMAL BORE DIAMETER 
VV(LB) PROJECTILE WEIGHT 
V(FPS) NORMAL MUZZLE VELOCITY 

O FIRING PROJECTILE FOR WHICH W/d s . 0.45 TO 0.55 
□ W IS UNKNOWN 



Figure 8. Approximate linear dependence of muzzle 
erosion rate on the square of the bore diameter for guns 
fired at three nominal muzzle velocity levels. (This 
figure has appeared as Figure 4 of NDRC Report No. 
A-357.) 


ferent calibers fired at about the same level of muzzle 
velocity with projectiles of normal weight for their 
caliber. (Such projectiles are characterized by having 
W/d z = .5, where W is projectile weight in pounds 
and d is bore diameter in inches.) This approximation 
is illustrated by Figure 8. 

10 4 6 Dependence on Muzzle Velocity 

Whether a gun erodes seriously at the muzzle 
seems to depend largely on its muzzle velocity. Quali¬ 
tatively it has been known for some years to ordnance 
personnel that guns of relatively low muzzle velocity 
(say, < 2,500 fps) are eroded at the muzzle only 


slightly, if at all, while guns of relatively high muzzle 
velocity (say, > 2,800 fps) are conspicuously eroded. 

It has been shown 67 that the quantity A d m /Nd 2 , 
(constant for guns of different calibers fired at about 
the same muzzle velocity) varies exponentially with 
the muzzle velocity, as represented by equation (2), 

A dJNd 2 = Ce v , (2) 

where V is the muzzle velocity, C is a constant, and e 
is the base of natural logarithms. A plot of log 
A d m /Nd 2 against V gave a straight line as shown in 



Figure 9. Muzzle erosion rate divided by square of the 
bore diameter (A dm/Nd?) for nonplated guns fired at 
different muzzle velocities. (This figure has appeared 
as Figure 6 of NDRC Report No. A-357.) 



2 

iS> 

o 


Ad m (IN) INCREASE IN AVERAGE BORE DIAMETER 
ACROSS LANDS AT MUZZLE 
-N NUMBER OF ROUNDS FIRED 
d(IN) NORMAL BORE DIAMETER 

W{LB) PROJECTILE WEIGHT 

V(FPS) NORMAL MUZZLE VELOCITY 

ol9 

o9 ./ 34 

o FIRING PROJECT 
W/d 5 • 0.45 TO 0.! 
□ W IS UNKNOWN 

ILES FOR WHICH 
55 

' 

270 2203124/< 

2o/ 

30° 25^°! 5 
s 

03 ”/ 

It 

> 

< 

01 , ^ 
020 ** 

'26 

x- olO 

>4 

□ 13 

29o o|2 
□ 06 

II 

08 

, , ■ 

_1_1_1_1_ 

-1-1-1-1- 

_1_ 


1500 2000 2500 3000 3500 


VELOCITY IN FPS 


Figure 10. Log (A dm/Nd?) for nonplated guns fired at 
different muzzle velocities. (This figure has appeared 
as Figure 5 of NDRC Report No. A-357.) 


Figure 10. Figure 9 is’the exponential form of the 
curve derived from the dashed line of Figure 10. Thus 
the early qualitative ideas regarding the effect of 
muzzle velocity on muzzle erosion are confirmed in a 
more quantitative way. 


CONFIDENTIAL 


























































MUZZLE EROSION 


207 


10,4 7 Correlation of Muzzle Erosion and 
Origin Erosion 

It should be emphasized that the foregoing descrip¬ 
tion of muzzle erosion and the correlations of muzzle 
erosion with the number of rounds, bore diameter, 
and muzzle velocity refer to the normal case in which 
origin erosion progresses simultaneously with muzzle 
erosion. What the rate of extension of muzzle erosion 
would be if origin erosion were prevented (as for ex¬ 
ample by the use of some erosion-resistant material 
or sufficiently cool powder) is difficult to predict. 

It has been suggested that muzzle erosion is 
“symptomatic” of origin erosion, owing to the pos¬ 
sible influence of the latter on the mode of behavior 
of the projectile in the bore. While this hypothesis 
seems reasonable, the information at hand 67 • 102 shows 
that muzzle erosion is not well correlated with origin 
erosion. The data of Table 6 illustrate the slight re- 


Table 6. Muzzle erosion of guns having the same degree 
of erosion at the origin of rifling, as indicated by meas¬ 
urements of average bore diameter with a star gauge. 67 



Gun 


Erosion, increase 
in diameter across 
lands (in.) 

At origin At 

Num¬ 
ber of 

Caliber 

Desig¬ 

nation 

Num¬ 

ber 

of rifling 
Ad 0 

muzzle 

A d m 

rounds 

fired 

155-mm 

Ml 

11 

0.152 

0.002 

1070 

155-mm 

Ml 

12 

0.112 

0.014 

1081 

155-mm 

M1A1 

409 

0.124 

0.041 

944 

8-in. 

Mk VI 

Mod 3A2 

190L2 

0.173 

0.024 

70 

8-in. 

Mk VI 

Mod 3A2 

173L2 

0.169 

0.054 

295 


lationship between muzzle erosion and origin erosion. 
The muzzle erosion of different guns of the same cal¬ 
iber varies widely when the extent of origin erosion 
is about equal. This means that factors other than 
origin erosion are also responsible for the growth of 
muzzle erosion. Therefore it would seem reasonable 
to suppose that the two kinds of erosion occur to 
some extent independently and that if origin erosion 
were eliminated the problem of muzzle erosion might 
still exist. 


10 4 8 Asymmetry of Muzzle Erosion 

Perhaps the most characteristic feature of muzzle 
erosion is its asymmetry, that is, the unevenness of 
the land erosion around the periphery at a given cross 


section and the tendency for the erosion to be con¬ 
fined largely to a group of adjacent lands. These 
effects are sometimes so great that they are readily 
noticeable on looking into the muzzle of a gun. 

Figure ll g shows the variation of land erosion 
around the periphery for three 14-in./45 cal. Naval 
guns, Mk VIII, at three stages of erosion. In these 
guns “pockets” of erosion have formed and the peak 
erosion is very nearly opposite the minimum erosion. 
Note that the peak erosion shifts in a clockwise direc¬ 
tion (viewed from the breech) with the number of 



6 9 12 3 6 9 12 

ORIENTATION (BREECH TIME) 


Figure 11. Asymmetrical muzzle erosion in 14-in./ 
45-cal. Naval guns, Mk VIII. (This figure has appeared 
as Figure 8 of NDRC Report No. A-357.) 


rounds fired. The curves are symmetrical and are 
very nearly like normal frequency distribution curves. 

These curves, however, are not to be taken as 
necessarily typical of 14-in. guns. Recent study of the 
muzzle of a 14-in. Naval gun after some 500 rounds 
showed two maxima and two minima in erosion 
rather than one maximum and one minimum as shown 
by the guns referred to in Figure 11. 130 

Figure 12 shows the variation of land erosion 
around the periphery for four 8-in. Army guns. These 

g In this figure and in Figure 12 breech time refers to posi¬ 
tions around the periphery, as viewed from the breech end. 
that correspond to those of a clock. 


CONFIDENTIAL 









































208 


DESCRIPTION OF ERODED GUN BORES 


curves are different from those of the 14-in. guns. 
Note the lack of symmetry of the curves and the ten¬ 
dencies toward formation of secondary peaks, most 
prominent in the Mk VI, gun Mod 3A2 at the 6 
o’clock position. The curve for the Ml gun, No. 20 
shows that the erosion is very nearly uniform around 
the periphery. Few instances of such uniform muzzle 
erosion have been found. 




> 



572 RDS 


•• # • ^ ' 


'"-^ 554 RDS 


Ml,NO.20 4IN. 




FROM MUZZLE 



FACE 

1 * 

1 


6 9 12 3 6 

ORIENTATION (BREECH TIME) 


Figure 12. Asymmetrical muzzle erosion in 8-in. 

Army guns. (This figure has appeared as Figure 10 of 

NDRC Report No. A-357.) 

It is not true, as is generally supposed, that the 
orientation of the peak erosion is always in a certain 
clock position or quadrant. Neither is it true that the 
lands included in the eroded sector are always those 
which are in the 6 o’clock position at the origin of 
rifling. However, Table 7 shows that the orientations 
of the peak erosion when referred to the origin of 
rifling are most frequently in the neighborhood of 6 
and 12 rather than 3 and 9 o’clock. This would seem 


to support the view that the position of asymmetric 
erosion is predetermined by the asymmetric position¬ 
ing of the projectile at the start of travel due to 
gravity. 

The asymmetry of muzzle erosion supports the the¬ 
ory that muzzle erosion is largely a result of mechan¬ 
ical action of the projectile as distinguished from 
powder gas erosion. As the powder gases do not spin 
they would not be expected to produce the preferen¬ 
tial erosion of a group of lands such as is usually 
observed in the muzzle end of a gun. Also, in 8-in. gun 
T2 No. 2, 8-in. gun Mk VI, Mod 3A2 and the 
14-in./50-cal. gun, the secondary erosion peaks are 
nearly opposite the primary ones, which would be 
expected from projectiles with the base bearing on 
one group of lands and the bourrelet bearing on op¬ 
posite lands. 

A comparison of the peaks for the 14-in. guns Mk 
VIII and for the 14-in. guns Mk VIII Mod A is in¬ 
structive. The peaks of the Mk VIII guns fall in the 
upper half of the bore circumference (9 to 3 o’clock) 
while those of the Mk VIII Mod A gun fall in the 
lower half (3 to 6 o’clock). The two types of guns 
differ in nominal land height (which would not be 
expected to determine orientation of erosion) and 
rifling twist. However, the eroded lands at the origin 
of rifling for both types of guns are in the lower half 
of the bore (3 and 9 o’clock) at the origin of rifling. 
This fact suggests that the rifling twist is one factor 
that governs the orientation of erosion and adds to 
the evidence that mechanical action of the projectile 
is of chief importance in muzzle erosion. 

10 4 9 Effect of Muzzle Erosion on 
Gun Performance 

Owing to the fact that origin erosion exists along 
with muzzle erosion, it is difficult to separate the 
effects of the two kinds of erosion on gun perfor¬ 
mance. An attempt was made to effect such a separa¬ 
tion by analysis of erosion and range data at the 
Naval Proving Ground but the data were found to be 
inadequate. 102 However, a general statistical method 
of attacking the problem was formulated which 
should prove useful if a considerable amount of ade¬ 
quate data is made available. 

We may, however, state some of the probable ef¬ 
fects of muzzle erosion on gun life. The maximum 
angle of yaw of a projectile is proportional to the yaw 
in the gun due to clearance between the projectile and 
the bore. 509 As muzzle erosion usually develops over a 


CONFIDENTIAL 



































MUZZLE EROSION 


209 


Table 7. Orientation of asymmetric muzzle erosion in ten major-caliber guns. Compiled from measurements of land 
height near the muzzle. 67 


Gun 

Rifling 

twist 

(turns/cal.) 

Approx, 
length 
of bore 
(in.) 

No. of turns 
from muzzle 
to origin of 
rifling 

Number 
of rounds 

Orientation* 
of “peak” 
muzzle 
erosion 

Orientation* 
at origin of 
rifling of 
“peak” muzzle 
erosion 

8-in., T2, No. 1 

1/25 

328 

1.64 

109 

2:00 

6:30f 

8-in., T2, No. 2 

1/25 

328 

1.64 

190 

2:30 

7:00f 

8-in. Mk VI, 

1/25 

289 

1.44 

159 

12:00 

6:30f 

Mod 3A2, 







No. 196L2 







14-in./45-cal. 

1/32 

533 

1.19 

35 E.S.R. 

11:30 

9:00$ 

Mk VIII 




70 E.S.R. 

12:30 

10:00 

No. 45L 




95 E.S.R. 

12:00 

9:30 

14-in./45-cal. 

1/32 

533 

1.19 

35 E.S.R. 

2:00 

11:30$ 

Mk VIII 




70 E.S.R. 

3:30 

1:00 

No. 65L 




95 E.S.R. 

4:00 

1:30 

14-in./45-cal. 

1/32 

533 

1.19 

35 E.S.R. 

8:30 

11:00$ 

Mk VIII 




70 E.S.R. 

10:00 

12:30 

No. 66L 




95 E.S.R. 

10:00 

12:30 

14-in./45-cal. 

1/25 

533 

1.52 

23 E.S.R. 

4:30 

10:30 

Mk VIII Mod A 




63 E.S.R. 

No peak 


No. 22L2 




103 E.S.R. 

6:30 

12:30 





128 E.S.R. 

6:30 

12:30 

14-in./45-cal. 

1/25 

533 

1.52 

23 E.S.R. 

3:00 

9:00 

Mk VIII Mod A 




63 E.S.R. 

5:00 

11:00 

No. 13L2 




103 E.S.R. 

6:00 

12:00 





128 E.S.R. 

6:00 

12:00 

14-in./45-cal. 

1/25 

533 

1.52 

25 E.S.R. 

No peak 


Mk VIII Mod A 




65 E.S.R. 

6:00 

12:00 

No. 15L3 




104 E.S.R. 

7:30 

1:30 





129 E.S.R. 

7:30 

1:30 

14-in. /50-cal.§ 

1/25 

607 

1.74 

209 E.S.R, 

1:00 

4:00 

Mk G, No. 110L2 







16-in./45-cal. § 

1/25 

617 

1.54 

140 E.S.R. 

9:00 

2:30 

Mk VI, No. 202 




200 E.S.R. 

9:00 

2:30 


* Breech time, 
t See Figure 12. 
t See Figure 11. 

5 Chromium-plated. 


considerable length of the forward end of the bore, it 
is reasonable to expect that such erosion will affect 
the value of the first maximum yaw. 

The results of a life test at Aberdeen Proving 
Ground of a chromium-plated 155-mmgun (M1A1E1, 
No. 3052), with a nonplated 155-mm gun (M1A1, 
No. 1069) as a control, seem to disclose the effect of 
muzzle erosion on yaw and yaw dispersion. Some 
2,000 rounds were fired in each gun, each round with 
the HE projectile, M101, and with a charge of single 
base propellant calculated to give a muzzle velocity 
of about 2,800 fps in a new gun. 

For the first 1,000 rounds the erosion of the chro¬ 
mium-plated gun at every measured position in the re¬ 
gion of the origin of rifling was less than that of 
corresponding positions in the nonplated gun. How¬ 
ever, in the neighborhood of the muzzle, the erosion 


of the chromium-plated gun was considerably greater 
than that of the nonplated gun. Measurements of 
yaw showed a consistently greater yaw for the chro¬ 
mium-plated gun than for the nonplated gun. Moreover 
the dispersion in yaw and the probable error in range 
were greater. These results (except for the probable 
error in range) are summarized in Figure 13. 

Muzzle erosion probably also increases the disper¬ 
sion in range because of possible leakage of gas past 
the projectile as it enters the enlarged area. A theory 
involving the following conditions in the bore has 
been proposed. 11 


h Personal communication from Dr. L. T. E. Thompson, 
Director of Research, Development and Test Organization, 
U. S. Naval Ordnance Test Station (Inyokern, California); 
formerly Chief Physicist, U. S. Naval Proving Ground (Dahl- 
gren, Virginia). 


CONFIDENTIAL 







210 


DESCRIPTION OF ERODED GUN BORES 


6.42 
6.38 
6.3 4 
6.30 
6.26 
6.22 
6.1 8 
6.1 4 
6.1 0 


EROSION AT THE 
ORIGIN OF RIFLING. 


M1A1ELP 


STAR GAUGE DIAMETERS 

IN A VERTICAL POSITION t 


.OM1A1 

'AT THE ORIGIN OF 
RIFLING. 






/ "J 

r 





NOTE: THE SCALE OF 
THE ORDINATE 
HERE IS TWICE 
THAT OF THE 
ORDINATE 
GIVING MUZZLE 
EROSION. 

I _ 

> 



// 








ROUNDS FIRED 


i<n 
2 u 


a: £ 
u < 
> > 



YAW 


— u - 

^MJAIEI 




' H— 

„ □ 


a 

r "C)__ 

___Q-- 

--MIA1 


8 






MIDDLE ROUND NUMBER 

Figure 13. Erosion and yaw for 155-mm guns, M1A1, 
No. 1069 (nonplated) and M1A1E1, No. 3052 (chromi¬ 
um plated). (This figure has appeared as Figure 2 of 
NDRC Report No. A-357.) 


The motion of the projectile is unstable in the pres¬ 
ence of high-pressure loading on the base. The axis of 
the projectile and the axis of the bore do not coincide 
because of clearance between the projectile and the 
bore. Since the rotating band is engraved by and fits 
closely the bore surface, there may be a point near 
the base of the projectile through which the instan¬ 
taneous longitudinal axis of the projectile always pass¬ 
es (approximately). In general, and particularly in a 
worn bore, the longitudinal axis and the axis of angu¬ 
lar momentum will not coincide and a sort of Poinsot 
motion may be expected. 

Under these circumstances, the bourrelet may roll 
in a spiral motion with considerable contact pressure 
along one section of the lands as the projectile moves 
toward the muzzle. When the projectile reaches the 
enlarged muzzle, the powder gases may escape past it 
in asymmetric fashion. As the projectile leaves the 
muzzle, the gas stream enveloping the projectile can 
be sufficiently asymmetric with respect to the projec¬ 
tile to introduce moments tending either to increase 


or decrease the yaw. If they tend to decrease the yaw, 
the flight of the projectile may be good. If they tend 
to increase it, a bad flight may be set up and it may 
last over a considerable portion of the early part of 
the trajectory depending on the extent to which 
damping is effective. According to this theory, when 
shots are fired successively from a gun with a badly 
eroded muzzle, some will have good flight, others bad 
flight. The result is a dispersion in range. 

Recent studies at the Naval Proving Ground of 
range and erosion data failed to establish muzzle ero¬ 
sion as affecting range dispersion. On the other hand 
they did not demonstrate that muzzle erosion was un¬ 
important. 102 

It has been suggested that in addition to the effect 
of the enlargement of yaw and the effect of the muz¬ 
zle blast in producing dispersion of yaw, the enlarged 
muzzle probably gives rise to an indeterminate angle 
and direction of departure of the center of gravity 
from the axis of the bore. Some evidence to support 
this suggestion has been adduced. 

Two experiments have been suggested to gain more 
exact information concerning the importance of muz¬ 
zle erosion apart from origin erosion: (a) that the 
performance of a new gun be compared to one in 
which the muzzle has been artificially eroded or to 
one in which the muzzle has been eroded by firing 
(the comparison to be made with the origin of rifling 
restored with a replaceable liner such as described in 
Section 26.3); and (b) that the performance be com¬ 
pared of two eroded guns having about the same bore 
contour at the origin of rifling, the one having con¬ 
siderable muzzle erosion produced in firing, the other 
having no muzzle erosion. The possibility of having 
the latter experiment (together with other controlled 
experiments to obtain further information on muzzle 
erosion) carried out by the Navy Department has 
been suggested by Division 1 to the Bureau of Ord¬ 
nance. 

10.4.10 ]VJ uzz l e Erosion and Body Engraving 

A phenomenon that may be associated with muzzle 
erosion is body engraving. In some guns and under 
certain conditions the recovered projectiles are found 
to be more or less deeply marked by the rifling at the 
base above and below the rotating band. In practi¬ 
cally all guns there is light marking of the bourrelet. 
Available information on the subject of body engrav¬ 
ing has been summarized in a recent NDRC report. 67 
In Appendix B of this report 67 appears a translation of 


CONFIDENTIAL 








































MUZZLE EROSION 


211 


what was perhaps the first summary of information 
on this subject: “Note sur les empreintes des projec¬ 
tiles au tir,” by P. Regnauld, Chief Engineer of 
French Naval Artillery, 1923. 

It has not been shown conclusively that body en¬ 
graving occurs in the muzzle, but available evidence 
is in agreement with such an assumption. In any case 
it appears that both body engraving and muzzle ero¬ 
sion are symptomatic of a common malady, namely a 
motion of the projectile in the bore which ultimately 
results in balloting or rubbing in the muzzle end or 
both. These two phenomena cause both an erosion of 
the muzzle and the marking of the projectile (body 
engraving). Figure 14 is an attempt to show what 
may be the separate effects of body engraving and 
muzzle erosion on the performance of the gun. It 
would seem that the only effect of body engraving per 
se would be the increased drag (and thus the loss of 
range) occasioned by the marking. The importance 
of this effect does not seem to have been evaluated. 

The effect of muzzle erosion is probably to increase 
the yaw and the dispersion of yaw (loss of range and 
dispersion of range) indirectly through leakage of gas 
according to the theory of Thompson (Section 10.4.9) 
and to increase the dispersion in range and lateral 
dispersion by a direct geometric effect, i.e., the effect 
of the enlargement itself on the angle of departure of 
the center of gravity of the projectile. 


If Figure 14 represents the relation between body 
engraving and muzzle erosion, it is clear that apart 
from correcting the behavior of the projectile in the 
bore, the correcting of body engraving may not cor¬ 
rect muzzle erosion and vice versa. Thus if we should 
protect the projectile against body engraving by use 
of a harder steel for the projectile, we might reduce 
body engraving, but there would still be wear or de¬ 
formation of the lands by the projectile and the muz¬ 
zle erosion would give rise to the effects suggested by 
Figure 14. On the other hand, if we should protect the 
bore against erosion, as for example by a satisfactory 
liner or coating, we might solve the problem of ero¬ 
sion ; but if the behavior of the projectile in the bore is 
bad, the projectile will still be body-engraved and 
perhaps emerge with a significant yaw. Therefore, as 
Figure 14 suggests, perhaps a consideration of the 
more fundamental problems of the motion of the pro¬ 
jectile in the bore 192 will lead to a solution of both the 
problems of body engraving and muzzle erosion. 

In Figure 14 the various factors that are shown as 
contributing to the motion of the projectile in the 
bore and to muzzle erosion are only in the nature of 
suggestions. The importance of each has been dis¬ 
cussed elsewhere. 67 The evidence presented there indi¬ 
cates that muzzle erosion is related more to the motion 
of the projectile than to the powder gases, abrasion 
by powder grains and other particles, and oxidation 


ORIGIN 



Figure 14. Suggestions as to the influence of the motion of the projectile in the bore on body engraving and muzzle 
erosion and the effects of these two phenomena on gun performance. 


CONFIDENTIAL 


















212 


DESCRIPTION OF ERODED GUN BORES 


by air after expulsion of the projectile, although the 
latter effects may contribute to some extent. 


10 5 NATURE OF THE ERODED SURFACE 

We have described the general erosion of guns in 
terms of the changes in the dimensions of the bore. 
We now turn to a detailed description of the eroded 
surface such as would be gained by examination with 
the naked eye or with the microscope. 

10,51 Cracking 

Perhaps the first thing that is noticed in examining 
the bore surface of a worn gun is the crack pattern, 
an example of which is given in Figure 15. Cracking 
always accompanies erosion in steel guns; the extent 
of their interrelation, however, is not clear. This sub¬ 
ject and the general one of the causes of cracking are 
discussed at length in Section 13.5. 

In advanced stages of wear, it is usually found that 
the widest and deepest cracks are longitudinal in the 
grooves and transverse on the lands. Circumferen¬ 
tially, cracks are usually deepest where lands meet 
grooves on the driving side. Longitudinally, they be¬ 
come shallower toward the muzzle. The depth of 
cracks varies considerably in different guns, but with¬ 


out any apparent regularity. A technique has been 
developed 98 whereby the three dimensional aspect of 
the crack system of an eroded bore specimen may be 
revealed by making a cast of the bore surface, as de¬ 
scribed in Section 11.4.4. 

The few laboratory studies that have been made of 
the muzzle ends of guns indicate that the depth of 
cracking bears no relation to the variation of erosion 
along the bore in the muzzle end. Thus, in 14-in. Na¬ 
val gun, Mk VII, Mod 1, No. 188L (A d m = 0.16 in.) 
cracking diminished while erosion continuously in¬ 
creased toward the muzzle. 130 

Development of Crack Patterns 

Microscopic studies of a number of worn guns of 
different calibers 49 ’ 112 have indicated that the devel¬ 
opment of the characteristic crack pattern near the 
origin of rifling in guns is about as follows. The initial 
pattern is largely controlled by the machine marks of 
the surface, which usually run transversely on the 
lands and longitudinally in the grooves. These areas 
of disturbed metal etch and crack minutely, with paral¬ 
lel grooves joined crosswise by little cracks, forming 
a brickwork pattern. As firing progresses, the little 
areas assume a roughly domed surface and become 
nearly equal in size while some of the cracks, in be¬ 
coming wider and deeper, form irregular shaped poly- 



Figure 15. Section of the bore surface at the origin of rifling of a 37-mm Browning automatic gun, Ml, after 8,032 
rounds. (This figure has appeared as Figure 5 of NDRC Report No. A-91, for which use it had been furnished by 
Watertown Arsenal.) 


CONFIDENTIAL 




NATURE OF THE ERODED SURFACE 


213 



gons bounding the smaller units. In a still later stage, 
distinct zigzag cracks more or less following the rifling 
are developed. Often the edges of the transverse 
cracks are modified (the side of the transverse crack 
toward the muzzle tending to be the lower). 

The development of the crack pattern and erosion 
has also been studied for chromium-plated guns. 86 In 
chromium-plated guns, firing weakens the bore by 
opening microscopic cracks originally present when 
the plate was deposited, by producing and developing 
new cracks such as those that form in steel barrels, 
and by effecting corrosion of the steel thus exposed. 
The sequence of events which leads to failure of a 
plated gun is given in Section 20.2.1. 

10 5 2 Liquefaction of the Bore Surface 

Examination of worn guns fired with single- and 
double-base powders has revealed so much evidence 
of liquefaction* of the surface during firing that there 
is little doubt that this phenomenon is one of the most 
important factors in the erosion of guns as they are 
fired today. This evidence appears under the stereo¬ 
scopic microscope as “ripples,” “tongues,” and smears 
of metal, gouges, and other irregularities. 112 Metallo- 
graphic examination of bore cross sections gives sup¬ 
porting evidence, summarized in Section 12.6, to liq¬ 
uefaction phenomena. 112 ’ 124 Melting of a number of 
nonferrous alloys used experimentally for gun bores 
has also been observed. 124 

In the muzzle section of a gun there is very little 
evidence of liquefaction except of the coppering. The 
erosion of this area seems to be the result chiefly of 
abrasion. 130 

The liquefied material may remain in place, may be 
deposited elsewhere, or may be swept out of the bar¬ 
rel. Whether it remains in place or is transported de¬ 
pends upon its mobility and the forces exerted upon 
it. A “pebbled” appearance is characteristic of a sur¬ 
face where the partially liquefied metal has solidified 
in place; tongues which stand out clearly in relief 
consist of material that had flowed as a “mush;” and 
relatively smooth, rippled surfaces, which may be 
covered with slender tongues of low relief, are evi¬ 
dence that melting had produced a liquid of high mo¬ 
bility. The last, although sometimes observed in guns 
fired with single-base powder (Section 12.5.3), is char¬ 
acteristic of guns that have been fired with double- 

5 The term "liquefaction” is preferable to "melting” since it 
appears that in many cases the melting that occurs is not of 
the original steel but of chemically altered steel. 


Figure 16. Pebbling in chamber of 5-in./51-cal. Naval 
gun liner, No. 806L2: Sirtgle-base powder; (a) 10X, 
(b) 35X. [Figure 19 (a,c) of NDRC Report A-440.] 


CONFIDENTIAL 










214 


DESCRIPTION OF ERODED GUN BORES 





Figure 17. Wind-rippled surfaces on which the tops of some ripples were blown forward into tongues and splashes 
(76-mm Army gun M1A1, Tube No. 1425; single-base powder), (a) 143^-in. from origin of rifling, 1000X, (b) 6)^ in. 
from origin of rifling, 250X , (c) U]/ 2 in. from origin of rifling, 500X. [Figure 9 (a, b, c), NDRC Report A-440.] 


CONFIDENTIAL 




215 


NATURE OF THE ERODED SURFACE 


base powders. Liquefaction resulting from the firing 
of these two different types of powder is discussed 
more fully in Section 13.2.4. 

Pebbling 

In the forward part of the chamber of a bag gun, 
where the surface is most exposed to the sweep of the 
powder gases, a brickwork crack pattern develops, 
such a pattern being started along the lines of the 
machine marks. The heat of the powder gases softens 
the metal in each little area bounded by cracks, and 
reactions with the powder gases (Section 13.3) lower 
the fusion range so that incipient liquefaction may 
occur. The metal tends to draw itself into a globule, 
with the result that a pebblelike pattern develops, as 
shown in Figure 16. In the chamber, the bulk-velocity 
of the propellant gases is not high and there is no rub¬ 
bing by the projectile; hence molten and previously 
molten material tends to remain in place. 

Forward of the origin of rifling, the pebblelike pat¬ 
tern may be somewhat obscured by the deposition of 
transported material over the cracks, in which case 
the rounding of the edges of the cracks in the original 
steel surface by liquefaction can be observed better in 
cross section. Metallographic examination has re¬ 
vealed that the inner white layer, which, as is de¬ 
scribed in Section 12.1.2, follows the contour of the 
steel around the edges of the cracks, is probably ma¬ 
terial that had been liquefied but not transported. 

Removal and Deposition of Liquefied 
Material 

Striking evidence of deposition of molten material 
has been found in a variety of guns. These deposits 
appear as splashes, tongues, and smears, often “but¬ 
tered” over the cracks, on the surface and mixed in 
the coppering as sheets, scales, and pellets. Examples 
of such deposits are shown in Figures 17 and 18. In 
the case of a “dirty” gun steel, an inclusion is often 
melted and “exploded out” in the direction of the gas 
stream, leaving a pit and becoming quenched as 
branched tongues just ahead of the pit, as shown 
in Figure 19. Sometimes streamlining “trails” are 
formed by crystals that did not go into solution in 
the partially liquefied metal. 

When liquefied material is transported, the direc¬ 
tion is not always axially toward the muzzle but may 
be affected by localized, turbulent flow of the gases. 
A discussion of deposits which show this is given in 



B 


Figure 18. Examples of smeared metal: (A) 3-in gun 
liner No. 1460, 50X; (B) Steel liner in a caliber .50 
heavy-barrel machine gun, 75X. (These figures have 
appeared as Figures 10(a) and 3(a) of NDRC Report 
No. A-440.) 

Section 13.4.1. Gouges and pits, direct evidence of 
the removal of material, are features that are pro¬ 
nounced also because of localized conditions. As was 
mentioned above, pits are left by the fluxing of inclu¬ 
sions; gouges, shown in Figure 20, are evidence of 


CONFIDENTIAL 





216 


DESCRIPTION OF ERODED GUN BORES 



scoring produced by localized gas leakage, which is 
discussed more fully in Section 13.4.1. The deposits 
ahead of gouges appear to have been sprayed from a 
jet. 

10 5 3 Distortion of Rifling 

Careful microscope, horoscope, and star gauge ex¬ 
aminations of a number of chromium-plated liners in 
caliber .50 machine gun barrels have given a clear 


grooves. The forward movement resulted in a piling 
up of steel forward of the enlarged area. The lateral 
movement was a flattening of the lands, in some cases 
widening them by as much as 50 per cent. Polished 
cross sections across the lands showed that the lands 
had spread out at the top, overthrusting onto the 
grooves, the top layer of steel carrying the chromium 
with it. The downward movement of the lands caused 


Figure 20. An example of gouging from blow-by in 
20-mm gun, Ml, No. 18709 fired with single-base pow¬ 
der, 6X. (This figure has appeared as Figure 6(a) of 
NDRC Report No. A-440.) 

a decrease in the diameter across the grooves. Figure 
21 is a sketch showing these effects of swaging of the 
rifling. 

Although the importance of swaging of the lands 
was first appreciated in the examination of chromium 
plated machine gun barrels, later evidence has indi¬ 
cated that swaging plays a part in the erosion of 
large-caliber chromium-plated guns. It is now be¬ 
lieved to be a general phenomenon, the result of 
which is frequently obscured in nonplated guns by 


Figure 19. Erupted surface with tongues of metal 
carried forward from pits left by the fluxing of inclu¬ 
sions in barrel fired with double-base powder in caliber 
.50 erosion-testing gun, 150X. (This figure has appeared 
as Figure 7(d) of NDRC Report No. A-440.) 


picture of how such barrels erode when fired with ball 
ammunition in moderately long bursts. 50 ’ 85 Although 
the chromium that was applied to the barrels offered 
remarkable protection from the action of the powder 
gases, the diameter across the lands at the origin of 
rifling increased almost as rapidly as if no chromium 
had been present. Owing to the mechanical action of 
the bullet the steel of the lands during firing had 
moved in three directions: (1) forward, (2) laterally, 
and (3) downward with lateral movement into the 


CONFIDENTIAL 







NATURE OF THE ERODED SURFACE 


217 


the fact that metal is removed simultaneously by 
powder-gas erosion. 

The swaging of the lands as related to the austen¬ 
ization of steel bore surfaces is discussed in Section 
13.4.2. The recognition of the importance of this ther¬ 
mal transformation as a factor in erosion has had 
considerable influence in orienting the search for 
erosion-resistant materials described in Chapter 16. 

In some guns firing long bursts at a rapid rate the 
bore is subject to heating in such a manner that en¬ 
largements and constrictions associated with differ¬ 
ential plastic deformations develop in the barrel. In 
extreme cases ricocheting bullets batter the bore and 
even perforate the barrel, as shown in Figure 16 of 
Chapter 5. 

10 5 4 Coppering 

Coppering is the term used to describe the deposi¬ 
tion of material from the copper or gilding metal ro¬ 
tating bands or jackets of projectiles onto the bore 
surfaces of guns. It varies in an irregular manner 
from round to round. It is commonly greatest in the 
central portion of the bore. The amount is usually 
insignificant at the origin of rifling unless the firing 
conditions have been relatively mild. Only in the case 
of machine guns which have been subjected to severe 
conditions is extensive coppering found at the muzzle. 
In a given gun the copper deposit is usually thicker 
in the grooves than on the lands. 

As was mentioned in Section 6.1.2, it was found in 
studies of bore friction 72 that the temperature of the 
interface between a bullet and the bore surface was 
above the melting point of copper, which indicates 
that thin films of the band metal are applied to the 
bore in a fused state. Thick deposits of copper are 
probably not remelted during successive rounds, but 
their configuration may be easily changed by the 
swaging action of the projectile. In the case of a 14-in. 
gun, a dark, spongy variety of copper appeared to 
have been sprayed into place beyond the region of 
maximum coppering and to have been slightly oxi¬ 
dized in the process. 98 

No coppering is found where general melting of the 
bore surface has taken place. In fact, it appears that 
relatively thick deposits cannot be built up when the 
bore-surface temperatures are much higher than the 
melting point of copper or gilding metal. Qualitative 
estimates were made of the amount and distribution 
of coppering relative to the bore-surface temperatures 
of the caliber .30 barrels used in the investigation 


described in Section 14.2. They showed that, with an 
increase in the bore-surface temperature during firing, 
there is a consistent tendency for the coppering to be 
displaced toward the muzzle. 99 Copper has been ob¬ 
served to have dripped out of the muzzle end of a 
caliber .50 machine gun barrel during the firing of a 
severe schedule. 

Relatively thick deposits of copper not only ob¬ 
scure the crack systems in gun bores but are rooted in 
the cracks, a fact which was at one time used as a 



/ ^-METAL MOVED FORWARD 

I Llano flattened 

Lland widened 

Figure 21. The lands of a caliber .50 machine gun 
barrel move in three directions when they are swaged 
by the impact of the bullet. 

partial basis for formulating a theory of the cause of 
gun erosion. 476 

The presence of coppering is more extensive than a 
superficial examination might indicate, for its surface 
is frequently blackened by copper oxide. 16 Moreover, 
it is sometimes so thoroughly contaminated with de¬ 
bris, such as dark, powdery, erosion products and car¬ 
bonaceous matter, that it may be completely coated 
over with dark material. 98 The fact that it entraps 
such substances makes it possible to study erosion 
products that might otherwise have been blown out 
of the gun. The technique of segregating these com- 


CONFTDENTIAL 






218 


DESCRIPTION OF ERODED GUN BORES 


pounds by dissolving the copper is described in Sec¬ 
tion 11.4.3. 

Coppering is considered by the U.S. Navy to be so 
important in its effect on ballistics that it employs 
decoppering agents and also specifies liberal bore- 
bourrelet clearances. 515 One of the decoppering me¬ 


thods is to add lead foil to the powder charges. An¬ 
other involves the use of a solution containing chro¬ 
mic and sulfuric acids. 


j The details of this method were worked out for the 
Naval Gun Factory by the Geophysical Laboratory. 


CONFIDENTIAL 





Chapter 11 

LABORATORY METHODS OF STUDYING GUN EROSION 

By Lloyd E. Line , /r. a 


111 INTRODUCTION 

T he most important laboratory methods of study¬ 
ing erosion that have been used or developed in 
the course of the work of Division 1 may be divided 
broadly into two classes: (1) Methods developed and 
employed to study the fundamental causes of gun 
erosion or to isolate and study separate effects, such 
as the effect of heat alone or of a single gaseous com¬ 
ponent on the surface of gun steel; and (2) methods 
used to test different liner materials, projectiles, and 
powders for their possible use in guns. 

The surest and most direct way to test a given 
material for erosion resistance, as is brought out in 
Section 16.2.5, is to incorporate the material in a gun 
of the same caliber and design that calls for improve¬ 
ment in durability. Indeed, this test should be per¬ 
formed ultimately, but less costly and less time- 
consuming methods are necessary to determine the 
materials that are most likely to prove the least erod- 
ible, so that final proving in larger calibers will in¬ 
volve a few carefully chosen materials. 

The caliber .50 erosion-testing gun was developed 
to serve both purposes mentioned above. Insofar as 
possible, this gun, together with its projectiles, was 
designed as a scaled-do wm model of a cannon. It fired 
at velocities above 3,500 fps. Such a gun was found to 
have wide applicability. In addition to its usefulness 
for testing the erosion of steel barrels and of liners of 
many materials with various propellants, it was em¬ 
ployed for the development of pre-engraved projec¬ 
tiles (Chapters 27 and 31) and of the Fisa protector 
(Chapter 32), for studies of body engraving of projec¬ 
tiles, and for studies of the fundamental causes of gun 
erosion. This gun and its uses are described in Section 
11 . 2 . 1 . 

The caliber .50 machine gun of standard design has 
been used extensively for testing metals and alloys 
(particularly stellite and chromium plate) described 
in Part V. The chief object of these experiments was 
the improvement of the caliber .50 w r eapon itself with 


a Technical Aide, Division 1, NDRC. (Present address: 
Chemistry Department, University of Tennessee, Knoxville, 
Tennessee.) 


the results described in Part VI. A description of the 
method of testing is given in Section 11.2.2. 

Erosion vent plugs have been used to test both bore 
surface materials and propellants and to study the 
causes of erosion. Thus vents of circular cross section 
have been used for testing alloys and powders (Sec¬ 
tion 11.2.3), for studies of the erosivity of different 
gases heated by adiabatic compression (Section 
11.3.1), and for testing the effect of stress on erosion 
(Section 11.2.5). 

Two types of apparatus using vents of D-shaped 
cross section were developed to study the chemical 
changes and the cracking that takes place when gun 
steel is exposed to powder gases. These are described 
in Section 11.2.4. 

The methods mentioned so far are not novel in 
principle. Several new procedures have been devel¬ 
oped to study various phases of the erosion problem. 
For example, as described in Section 11.2.6, filings of 
metallic or other specimens were mixed with the pow r - 
der charge for a caliber .30 round. The rifle w'as fired 
into an evacuated glass tube. Some of the filings were 
recovered from a deposit in the walls of the glass tube 
and the nature of the alteration produced by the pow ¬ 
der gases was examined by x-ray diffraction. 

Although it turned out that cavitation erosion bore 
no relation to erosion in guns, some experiments were 
early carried out in which different metals and alloys 
were subjected to magnetostriction oscillation in air- 
free water. The apparatus is described in Section 
11.3.3. 

In an effort to evaluate the causative influence on 
erosion of the high bore-surface temperature apart 
from any chemical causes, a method was developed 
(Section 11.3.2) in which the surface of a small sample 
of gun steel or other material was bombarded by 
bursts of electrons. In this way a very thin layer at 
the surface of the specimen was subjected to a very 
high temperature for a period of time comparable to 
the time of passage of a projectile in a gun bore. 

In addition to these laboratory methods of produc¬ 
ing gun erosion or some particular feature of erosion, 
special techniques were developed to study the prod¬ 
ucts of erosion in guns and in the devices enumer¬ 
ated above. The methods used to disengage and to 


CONFIDENTIAL 


219 



220 


LABORATORY METHODS OF STUDYING GUN EROSION 


segregate products of erosion are described in Section 
11.4. The use of the methods of chemistry, metal¬ 
lography, and physics (x-ray and electron diffraction, 
radioactive tracers) in studying the products of ero¬ 
sion are discussed in Section 11.5. 

“ 2 LABORATORY FIRING TESTS 
11,2,1 Caliber .50 Erosion-Testing Gun 122 

Prior to 1942 most of the laboratory methods to 
determine the erosiveness of powders and the resis¬ 
tance to erosion of various materials were of the vent- 
plug type, which do not completely simulate gun con¬ 
ditions. Therefore in the early days of the work of 
Division 1 there seemed to be a need for (a) an erosion¬ 
testing gun which could be used to test materials un¬ 
der normal and hypervelocity conditions, (b) a hyper¬ 
velocity gun that would provide data translatable to 
guns of larger calibers, and (c) a standardized erosion 
test in which all resistant materials and propellants 
would be examined under identical conditions. 

Design of Gun 

With this end in view the caliber .50 erosion-testing 
gun was developed by the Franklin Institute for Di¬ 
vision 1, NDRC. This gun was composed of three ma¬ 
jor units (a) the receiver and mount, (b) the firing 
mechanism, and (c) the test barrel. 

The test barrel was a standard 45-in. monobloc 
barrel modified as shown in Figure 1. The chamber, 
origin of rifling, and bullet seat were shaped to receive 
a 20-mm cartridge case necked down to hold a caliber 
.50 bullet. The depth of rifling was made 0.010 in. in¬ 
stead of the usual 0.005 in. The enlarged chamber 


permitted a greater powder charge for attainment of 
high velocities and the increased depth of rifling pro¬ 
vided a greater area of contact between the rotating 
band of the bullet and the rifling, a necessary condi¬ 
tion because of the high velocities used. 

Of many of the materials tested, it would have 
been difficult to make a full-length barrel. Further¬ 
more, since erosion is confined to the region of the 
origin of rifling, this was unnecessary. Accordingly, a 
special gun-barrel assembly was developed into which 
short liners could be inserted. The assembly consisted 
of three units as shown in Figure 2: (a) the cham¬ 
ber section, (b) the breech, and (c) the muzzle section. 
The liner was 8 in. in length and was pressed into 
the breech. This gun was very economical because 
the various units could be used many times. 

Firing Conditions 

The firing conditions were chosen so as to impart a 
muzzle velocity of 3,500 to 3,750 fps to a 710-grain 
projectile fired from the 45-in. barrel. This required a 
powder charge of 476 grains of double-base (20% ni¬ 
troglycerin) powder of a web size that would keep the 
maximum pressure in the range 56,000 to 58,000 psi 
(copper). Occasionally a single-base (IMR) powder 
was employed to observe effects with a cooler powder 
at the same muzzle velocity. 

The resistance to erosion of different materials and 
the gun performance with those materials were stud¬ 
ied with three principal types of projectiles: (a) the 
caliber .50 ball bullet, M2, (b) a copper-banded artil¬ 
lery-type bullet, and (c) a steel pre-engraved projec¬ 
tile. The ball bullets were used with steel barrels of 
standard depth of rifling for erosion tests when the 
accuracy life was not of any interest. 



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LABORATORY FIRING TESTS 


221 



Figure 2 . Liner-barrel assembly for caliber .50 erosion-testing gun. (This figure has appeared as Figure 7 of NI)RC 
Report No. A-450.) 


The artillery-type bullets were fired at hyper¬ 
velocity to determine velocity-life and accuracy-life 
of a liner assembly. This bullet, shown in Figure 3A, 
was so designed that accuracy would not be too crit¬ 
ically affected by erosion at the origin of rifling and so 


as to produce minimum interference of the band - with 
the grooves. A band width of 0.50 in. was found to be 
necessary in order to yield sharp engraving without 
widening of the grooves. 

The steel pre-engraved projectile in Figure 3B, 




Figure 3. Artillery-type (A) and pre-engraved projectiles (B) for caliber .50 erosion-testing gun. (This figure has 
appeared as Figures 11 and 12 of NDRC Report No. A-450.) 


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222 


LABORATORY METHODS OF STUDYING GUN EROSION 


which is discussed in detail in Chapter 27, was 
used in special tests when it was desired to study the 
behavior of a liner or bore surface coating in the 
nearly complete absence of engraving forces and 
friction. 

Three firing schedules were standardized permit¬ 
ting a choice of the severity of the conditions. These 
schedules and their uses are given in Table 1. After 
each group of 35, 70, or 140 rounds, respectively, the 
gun was plug gauged and star gauged, examined with 
a boroscope, and the bore surface photographed. In 
most cases the liners or barrels were examined metal- 
lographically. 77 ’ 124 Control tests were made with gun 
steel barrels (SAE 4150, modified) for each of three 
firing schedules. 

Criteria for Barrel Failure 

Gun barrels or liner assemblies were considered to 
have failed when (a) velocity had dropped 200 fps, 
(b) the mean radius of dispersion at 100 ft had in¬ 
creased to three times its initial value or when bullets 
produced keyholes, or (c) the liner material had 
changed in some way that prevented further firing. 

Maximum powder gas pressures were measured 
with a crusher gauge of the copper cylinder type. Ve¬ 
locity was determined with two screens 37 ft apart 
connected to an Aberdeen chronograph, the first 
screen being 8 ft from the muzzle. The mean radius of 
dispersion was calculated, in many of the tests, from 
targets placed at 100 ft. 

Erosion on the lands was followed with a set of 
plug gauges which would give readings in steps of 
0.002 in. from 0.490 to 0.516. Erosion across the 
grooves could be measured with a set of rifled plug 
gauges by steps of 0.002 in. from 0.511 in. to 0.529 in. 


These measurements were plotted as bore profiles at 
various stages of erosion. These gauges were very sat¬ 
isfactory as far as simplicity and speed were con¬ 
cerned, but they could be used only in measuring uni¬ 
formly tapered erosion. They were not suitable, for 
example, in measuring the erosion of bore surfaces 
which had been plated. A small area of the resistant 
material adhering to the surface would easily stop the 
advance of the plug gauge. To avoid this difficulty a 
star gauge was constructed that could measure land 
and groove diameter within 0.0002 in. 

The progress of erosion was also followed by both 
visual and photographic examination of the bore sur¬ 
face through a boroscope. For many tests, barrel tem¬ 
peratures were ascertained by attaching iron-con- 
stantan thermocouples on the outside of the barrel at 
10K in- from the breech face. 

Projectiles were recovered in screened sawdust 
after each group of rounds. Measurements of band 
diameter, width and depth of engraving, and observ¬ 
ation of body engraving indicated the effect of erosion 
on the behavior of the projectile in the bore. 

Uses of Erosion-Testing Gun 

The caliber .50 erosion-testing gun had a wide 
range of usefulness as a research tool. Although pri¬ 
marily designed and most extensively used for the 
first of the problems listed below, it has been em¬ 
ployed to study many others. 

(1) Determination of the erosion-resistance of 
materials under hypervelocity conditions (Section 
16.3.5). 

(2) Determination of the erosiveness of various 
propellants at 3,300 fps (Chapter 15). 

(3) Design and behavior of projectiles for hyper- 


Table 1. Firing schedules used with caliber .50 erosion-testing gun. 


Schedule 

Cycle 

Use 

I 

(1) 10 rounds to determine pressure and velocity 

(2) 20 rounds at 4 rounds per min for erosion 

(3) 5 rounds for projectile recovery 

For preliminary test of a new material for erosion resistance, 
with ball bullets M2; rifling depth 0.005 in. 

II 

(1) 10 rounds to determine pressure and velocity 

(2) 55 rounds at 4 rounds per min for erosion 

(3) 5 rounds for projectile recovery 

Adopted as standard for barrels with rifling depth of 0.010 in. 
and PE projectiles, for testing erosiveness of powders and 
erosion resistance of liners and coatings. 

III 

(1) 10 rounds to determine pressure and velocity 

(2) 130 rounds at 6 rounds per min for erosion 

Adopted as a standard for testing liners and coatings more 
resistant to erosion than gun steel, using ball bullets, M2 
and a rifling depth of 0.010 in. 


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LABORATORY FIRING TESTS 


223 


velocity guns, particularly of pre-engraved projectiles 
(Chapters 27 and 31). 

(4) Design and behavior of the Fisa protector 
(Chapter 32). 

(5) The fundamental causes of erosion (Chapter 
13). # 

(6) Erosion at the joints of liners (Section 10.3.8); 
and 

(7) The causes of body engraving (Section 10.4.2). 

At about the same time that the caliber .50 erosion 

testing gun was planned, the Army Ordnance Depart¬ 
ment was interested in a device for accelerated ero¬ 
sion tests. What was termed an “erosion gauge” was 
developed by Frankford Arsenal 313 in cooperation 
with Aberdeen Proving Ground and Watertown Ar¬ 
senal. This device was similar in principle to the cal¬ 
iber .50 erosion-testing gun, but differed in that the 
bore was caliber .30. Experience in its use, 314 ’ 315 indi¬ 
cated that the smaller size offered no marked advan¬ 
tage and some disadvantages. Thereupon it was de¬ 
cided by the Army Ordnance Department that its 
establishments should use the type of erosion-testing 
gun developed by the Franklin Institute. 

112-2 Caliber .50 Machine Guns 

Introduction 

The hypervelocity caliber .50 erosion-testing gun 
just described had one serious limitation, which was 
that rapid-fire tests could not be made with it. b The 
only reliable machine guns available for such tests 
were the Browning caliber .30 and caliber .50. Pre¬ 
liminary trials with the caliber .30 gun (Section 
16.3.7) indicated that the rate of erosion was too 
slight, and therefore attention was concentrated on 
the caliber .50. At first it was hoped that by control 
of the barrel temperature of this gun through varia¬ 
tion of the rate of fire it would be possible to simulate 
the conditions of erosion in a gun of higher velocity 
and larger caliber, such as the 120-mm gun, Ml. For 
this purpose the special firing schedule (designated as 
“GL-135” in Section 23.1.3) was developed. 49 

Subsequent experience in the testing of materials 
by this and other schedules led to the conclusion that 
it is not possible to evaluate completely hypervelocity 
performance in this way. Nevertheless the caliber .50 

b It had been planned to overcome this deficiency by adapt¬ 
ing a 20-mm automatic gun mechanism to use the barrel of the 
caliber .50 erosion testing gun. Some preliminary tests were 
made with Model I of the mechanism described in Chapter 28 
shortly before Division l’s contracts were terminated. 


machine gun proved to be a useful tool for the study 
of erosion in two ways. The observations made on 
steel barrels, both plain and chromium plated, after 
they had been fired in this gun, formed the basis for 
some of the conclusions expressed in Chapter 13 con¬ 
cerning the causes of erosion. This was especially true 
with respect to the phenomenon of swaging of the 
rifling, described in Sections 10.5.3 and 13.4.2. 

In the second place, this gun served well for tests 
of the erosion resistance of metals and alloys, as out¬ 
lined in Section 16.3.8. For this purpose they were 
prepared either as short breech liners or as coatings 
on a steel barrel or steel liner. Eventually it became 
possible to improve the life of caliber .50 machine gun 
barrels by the use of stellite liners (described in Chap¬ 
ter 22), of tapered chromium plate applied on a 
hardened steel bore surface ( Chapter 23), and finally 
by a combination of these two improvements (de¬ 
scribed in Chapter 24). 

Geophysical Laboratory Liner Design 81 

The first type of liner assembly was developed at 
the Geophysical Laboratory, CIW in 1942 and 
1943, for testing the erosion resistance of a number of 
metals and alloys that appeared promising as a result 
of erosion vent plug tests (Section 11.2.3). The heavy 
barrel 0 was selected for this purpose because its large 
wall thickness near the breech end facilitated the in¬ 
sertion of a liner. 

The design of liner assembly shown in Figure 4 
proved to be the most satisfactory one for inserting a 
liner in a heavy barrel. The liner was inserted by 
pressing it into the middle piece on a taper. The other 
two pieces were screwed to the middle one, as shown, 
to form the completed assembly. The liner was usu¬ 
ally 5 in. long but a 9-in. liner was sometimes inserted. 

If the barrel was not subjected to too severe a firing 
schedule, the assembly could be used for another test 
by inserting a new liner. Although it was found that 
the press fit was sufficient to prevent the rotation of a 
liner under normal conditions, it was usually secured 
against rotation by a pin extending from the outside 
of the barrel into the liner to within J/f 6 in. from the 
bore. 

Since the joint at the rear end of the liner was cov¬ 
ered by the cartridge case, no special method of seal- 

c By “heavy barrel” is meant the 45-in. barrel weighing 28 
to 30 lb. Barrels made according to four different drawing 
numbers have been used: D28253-11, D28253A, D28253-A3, 
and D28269-8X. These differ only in small variations of out¬ 
side contour. 


CONFIDENTIAL 





224 


LABORATORY METHODS OF STUDYING GUN EROSION 



Figure 4. Design of liner for caliber .50 heavy machine gun barrel used at Geophysical Laboratory for test of materials. 
(This figure has appeared as Figure 2 of NDRC Report No. A-409.) 


ing was used there. In order to prevent the powder 
gases from leaking through the forward joint into the 
space between the liner and the barrel, a washer was 
inserted at this junction. This washer is not shown in 
Figure 4. 

It was desirable to use as thin a liner as possible, 
first, because it was very difficult to obtain thick rods 
of some of the materials, and second, in order to leave 
as great a thickness of gun steel as possible so as not 
to weaken the barrel unduly. A minimum wall thick¬ 
ness for the liner of x /% in. was decided upon following 
the failure of a liner made of caliber .50 gun steel 
which was but % i n - thick. 

As described in Part VI these liners were used to 
test electroplates of chromium and other metals, 
coatings of the bore surface, gun steel, special steels, 
molybdenum, Stellite No. 21, and other materials. It 
was necessary to modify the assembly described 
above in order to test certain materials, particularly 
molybdenum. 

Crane Company Liner Design 80 

The design of liner for the heavy barrel just de¬ 
scribed was also used by the Crane Company in its 
first experiments with testing erosion-resistant ma¬ 
terials. Then, when stellite was introduced, it was 
found necessary to develop a new design of liner hav¬ 
ing a flange, as described in Section 22.2. This design, 
which used a liner 9 in. long, shrink-fitted into a steel 
barrel and held in place by a steel retainer, is shown 
in Figure 1 of Chapter 22. 

The later success of this type of liner adapted to 
caliber .50 aircraft barrels d and its production in large 
quantities had the important advantage, as far as fur- 

d By “aircraft barrel’' is meant the 36-in. barrel weighing 
about 10 lb made according to Ordnance Department drawing 
D35348A or D28272. 


ther erosion investigations were concerned, that a 
source of barrels recessed to receive liners greatly 
simplified the testing of a wide variety of bore-surface 
materials with the results described in Section 16.4. 

A modification of the Crane design of liner, which 
differed in that it did not use a shoulder flange, was 
used for chromium-base alloys, as described in Sec¬ 
tion 17.4.3. 

Test Procedure 

The firing schedules to which both these liner as¬ 
semblies were subjected varied according to the infor¬ 
mation desired. For example, in comparing liners of 
metals showing high resistance to erosion, relatively 
severe schedules had to be employed. The various 
schedules that were used, and the applicability of 
each, are summarized in Section 23.1.3. 

The performance of the liner assemblies was judged 
by the muzzle velocity and accuracy. Other data, 
such as gauge measurements, hardness, microscopic 
observation, and some temperature measurements, 
were obtained in a routine manner. Ballistic data for 
caliber .50 ammunition have been summarized in a 
report 312 from Frankford Arsenal. 

11,2,3 The Erosion Vent Plug 

General Description 

Vent-plug tests have long been used by a large 
number of investigators as a means of studying gun 
erosion. A number of summaries 16 ’ 260t 479 of the vent- 
plug method have appeared. 

The explosion vessel that is commonly used for ero¬ 
sion vent-plug tests is arranged with three openings. 
One is fitted with some means of firing the charge, 
such as an electric connection for firing the powder by 


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LABORATORY FIRING TESTS 


225 


heating a wire to incandescence; a second opening is 
fitted with a pressure gauge; and the third is arranged 
so that a cylindrical test-plug with a small axial hole 
can be inserted in it. The gases formed when a charge 
of powder is fired in the explosion vessel escape 
through this hole with high velocity and enlarge its 
diameter by removal of metal. The extent of vent- 
plug erosion is usually determined by measuring the 
loss of weight of the plug. 

Improved Vent-plug Apparatus 

A vent-plug apparatus, employing less severe con¬ 
ditions than those used by earlier investigations has 
been developed. 27 The use of precise laboratory tech¬ 
niques contributed to the reproducibility of the re¬ 
sults, which was achieved through better control of 
the rate of burning of the powder by closing the muz¬ 
zle end of the vent with a rupture disk. The disk 
allowed the powder gases to be released always at the 
same predetermined pressure. 

The apparatus, shown in Figure 5, consisted of a 
cylindrical steel block having four holes that met at 
the center, into each of which was screwed a steel 
plug. When a charge of propellant was exploded inside 
the block, pressure was built up until a copper or 
brass disk ruptured and released the gas through the 
vent. In this apparatus there was provision for both a 
crusher gauge and a piezoelectric gauge. The erosion 
plug A was a cylinder in. in outside diameter and 
% in- long. 

Two sets of test conditions were adopted in testing 
with this apparatus: (1) A vent 3^6 in. in diameter 
subjected to the action of the gases from double-base 
powder (4.3 g, loading density 0.187 g/cc, which pro¬ 
duced maximum pressures in the neighborhood of 
35,000 psi with a new vent); (2) A vent 3^8 in. in di¬ 
ameter and a charge of single-base powder chosen to 
yield the same maximum pressure as above. Condi¬ 
tions (1) are more severe than those of (2). The latter 
give erosion rates for steel samples that are roughly 
equivalent to the rate of increase in groove diameter 
per round found in certain medium-caliber guns. 75 

The vent plug has been used for comparing the 
erosion of many metals and alloys both with single¬ 
base and double-base powders, as recorded in Chap¬ 
ter 16. The erosiveness of various propellants was 
compared by means of vent plugs of gun steel, as 
described in Chapter 15. 

There are several objections to using the data ob¬ 
tained from vent-plug tests, without further consider¬ 


ation, to evaluate the expected performance of a bar¬ 
rel or liner made of a new material. For instance, such 
tests do not indicate resistance to mechanical wear 
such as by abrasion and swaging. The question of the 
interpretation of the results of vent-plug tests from 
the viewpoint of gun erosion has been discussed else¬ 
where. 16 



Figure 5. Erosion vent-plug apparatus. The ignition 
plug K, which carried a platinum filament for electric 
ignition, was screwed into the horizontal hole N after 
the powder had been placed in that hole. (This figure has 
appeared as Figure 1 of NDRC Report No. A-148.) 

Rifle Barrel Used as a Vent 

The vent plug, in an apparatus very similar to that 
described above, was used to study erosion by mix¬ 
tures of carbon monoxide and carbon dioxide. 62 Car¬ 
bon monoxide and oxygen in various proportions 
were introduced under pressure into the explosion 
chamber and ignited electrically. A series of ratios of 
carbon monoxide to dioxide in the eroding gases was 
obtained by varying the oxygen content of the ex¬ 
ploding mixture. These tests and their results are dis¬ 
cussed in Section 14.3. 

One of the objects of this research was to ascertain 
the possible importance of the formation of iron car¬ 
bonyl in erosion. 63 In order to recover a detectable 
amount of iron carbonyl it was necessary to increase 
greatly the area of the surface exposed to the gases 
over that of the standard vent plug. This was accom¬ 
plished by using the bore of a B-17 Enfield rifle as a 


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226 


LABORATORY METHODS OF STUDYING GUN EROSION 



CONFIDENTIAL 


Figure 6. Apparatus used in the collection of the gaseous products of erosion of a rifle barrel adapted to vent-plug experiments. (This figure has 
appeared as Figure 31 of NDRC Report No. A-311.) 




























































































































































































LABORATORY FIRING TESTS 


227 


vent. The apparatus is shown in Figure 6. The explo¬ 
sive mixture of oxygen and carbon monoxide was 
contained in a pressure vessel of 62.5-ml volume. The 
rifled portion of an Enfield barrel was machined to 
thread directly into the pressure vessel. The breech 
end was closed with a rupture disk which would burst 
at a piezo pressure of 25,500 psi. A thinner rupture 
disk which would burst at 10,500 psi closed the muz¬ 
zle end. 

The muzzle end was machined for attachment, in a 
vacuum-tight manner, to a large sealed tube of copper 
called the muzzle tube. The first 9 in. of this tube 
were lined with glass. 

The muzzle tube was connected through a Kerotest 
valve to a glass trap cooled in a bath of liquid air. 
Volatile products of erosion were condensed in this 
trap. The original report 63 contains complete descrip¬ 
tions of the development of the above arrangement, 
of many pieces of auxiliary apparatus, such as that 
used to purify the carbon monoxide from iron penta- 
carbonyl, and of analytical apparatus for identifying 
the presence of volatile iron compounds, in particular 
the carbonyl, in the products condensed in the cooled 
trap. 

In performing the experiment the muzzle tube and 
trap were well evacuated. The Kerotest valve was 
closed momentarily while the mixture of carbon mon¬ 
oxide and oxygen, or in some experiments a standard 
solid propellant, was ignited. The resulting gases and 
vapors were then slowly pumped out through the 
trap. The trap was sealed off and removed about 30 
min later, and the volatile products were transported 
through a heated capillary tube with a carrier gas 
while the trap was slowly warmed. A deposit of dark- 
colored material in the fine capillary, illustrated in 
Figure 11 of Chapter 14, indicated the presence of a 
volatile iron compound. If, when the material was de¬ 
composed, analysis of the gaseous product indicated 
the presence of carbon monoxide, the volatile com¬ 
pound was taken to be iron carbonyl. 

The identification of iron carbonyl among the ero¬ 
sion products in these experiments with gas mixtures 
and solid propellants and a.description of the various 
solid deposits in different parts of the apparatus is 
fully described in Section 14.3.4. 

11,2,4 Erosion Vents with 

D-shaped Cross Section 

Studies of Cracking and Mild Erosion 

An apparatus similar in principle to that employed 


with the usual vent plug, described in Section 11.2.3, 
was devised 51 to study surface changes such as crack¬ 
ing. The surfaces were to be observed and the effects 
to be photographed with a comparison microscope. 
It was desirable to follow the course of surface changes 
through many rounds. The apparatus was designed, 
therefore, so that removable specimens with flat sur¬ 
faces could be used. The use of this method to eval¬ 
uate metals and alloys with respect to erosion resis¬ 
tance is mentioned in Section 16.3.1. 

Figure 7 shows a cross section of the apparatus. 
The “gun” G in which the charge was burned was the 
receiver of a caliber .30 Army rifle, M1903, and the 



Figure 7. Modified erosion vent apparatus (with 
D-shaped vents) for study of surface cracking. (This 
figure has appeared as Figure 1 of NDRC Report 
No. A-271.) 

charges were prepared in caliber .30 cartridge cases. 
The test specimen S was made from machined round 
stock. Two flat surfaces were milled on diametrically 
opposite sides of the rod which was inserted with a 
good fit into the hole in the cone-shaped plug P. The 
removable steel bridge B prevented the specimen 
from being ejected as a projectile from the explosion 
vessel. The apparatus could function without the col¬ 
lar C but this was found to prevent occasional loosen¬ 
ing of the cone-shaped plug in the wall of the vessel. 
Maximum pressure was controlled by a brass rupture 
disk R. When inserted, the specimen provided D- 
shaped vents through which the hot powder gases 
passed over the flat surfaces. The surfaces were of dif¬ 
ferent widths; hence the cross sections of the vents 
w^ere different and two sets of conditions could be 


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228 


LABORATORY METHODS OF STUDYING GUN EROSION 


studied simultaneously. The exposed wall of the hole 
in the plug P was, of course, also subjected to erosion. 
The plug P had to be replaced from time to time to 
keep the conditions in the vents constant. 

The amount and kind of powder used, and the 
thickness of the rupture disk controlled the time- 
temperature relation, the composition of the eroding- 
gases, the maximum pressure, and the rate of rise of 
pressure. All these things were important in obtaining 
the desired conditions. Because the D-shaped orifices 
had a much smaller cross-sectional area than the hole 
punctured in the rupture disk most of the gas went 
out through the rupture disk when it broke. The frac¬ 
tion of the total gas that passed over the test surfaces 
depended on the rate of rise of pressure before the 
rupture disk broke and the maximum pressure which 
was attained at the moment when the disk ruptured. 
With this arrangement it was possible to erode the 
test surfaces at rates much more comparable with 
erosion rates in guns than is practicable with the con¬ 
ventional vent plug. Erosion rates with this apparatus 
are plotted in Figure 13 of Chapter 14. 

In order to study cracking, it was necessary to se¬ 
lect a powder and size of charge that would not cause 
surface melting but would be severe enough to exhibit 
cracking after a few rounds. It was found that the 
conditions with 2.5 g of caliber .45 pistol powder and 
a 346-im rupture disk caused noticeable cracking of 
SAE 4140 gun steel in four rounds; these conditions 
were chosen as standard. In order to rate the differ¬ 
ent materials, the procedure was to fire a certain 
number of rounds and compare the degree of cracking. 
The results obtained with a variety of metals and 
alloys are reported elsewhere. 51 

Although the conditions of temperature and pres¬ 
sure as a function of time and the composition of the 
eroding gases were quite similar to those encountered 
in guns there were two departures from gun-firing 
conditions which may be important in the study of 
cracking. To some extent, the powder gases streamed 
over the test surfaces during the period in which the 
pressure was building up, whereas in guns the gases 
come over as a single blast when the projectile passes. 
This probably does not constitute a radical departure 
because in a similar vessel it could be shown that, 
with a powder burning as quickly as pistol powder, 
the major portion of the gas issuing through the D- 
vents does so after maximum pressure is attained. 

A more radical departure from the conditions in a 
gun occurs because of the lack of any wiping action. 
Thus a coating of the products of a reaction between 


the specimen and the powder gases builds up at the 
test surface. This coating is not subject to removal as 
it would be in a gun by the passage of a projectile. A 
film of thermal insulation may thus be established 
when several rounds are ignited in succession, and, in 
fact, a difference in the degree of cracking was ob¬ 
served between a specimen cleaned off after every few 
rounds and one subjected to from 20 to 30 rounds 
without removing the reaction products. 

The apparatus described above was also used in 
the study by a radioactive tracer technique, discussed 
in Section 14.4, of the effect of sulfur and other com¬ 
ponents of black powder on the erosion of gun steel. 
The flat surfaces of the test rods facilitated the beta- 
ray measurements. 53 

Apparatus Used in Studies with X-Ray and 
Electron Diffraction 

Much information concerning the chemical alter¬ 
ation of surfaces exposed to powder gases was ob¬ 
tained by means of x-ray and electron diffraction 
(Section 11.5.2). For some of this work an explosion 
vessel, shown in Figure 8, was designed in which the 
pressure and temperature as a function of time cor¬ 
responded approximately to those in a medium- 
caliber gun. 31 This method, like the one described 
above, was also used to a limited extent to test ma¬ 
terials (given in Section 16.3.1) for erosion resistance. 

Following ignition of the powder, the gases passed 
over two sides of a rectangular specimen of gun steel 
(0.1x0.2x0.8 in.), held in a cylindrical opening, and 
escaped from the vessel through a vent plug. The size 
of the orifice in the vent plug, usually Xe in- in di¬ 
ameter, and the amount and type of powder were the 
factors which determined the pressure in the explo¬ 
sion chamber. With 3.5 and 2.5 g FNH-M1 powder 
the measured maximum pressure was 51,000 and 
37,000 psi respectively; with 2.5 g of a specially pre¬ 
pared double-base powder the maximum pressure was 
43,000 psi. 

Scrupulous care was taken to prevent contamina¬ 
tion of the test rods with the result that no foreign 
material, particularly inorganic salts, was detectable 
in the diffraction patterns obtained from the surfaces 
of control specimens. 

In such studies it is naturally necessary to avoid 
melting, but it is also necessary to obtain a certain 
minimum thickness of reaction products even though 
under ideal conditions it is possible to obtain good 
electron diffraction patterns with layers only a few 


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LABORATORY FIRING TESTS 


229 




PRESSURE GAUGE 



IGNITER 



SECTION AB 


note: all dimensions are in inches 

Figure 8. Second explosion vessel using D-shaped vents. (This figure has appeared as Figure 1 of NDRC Report 
No. A-199.) 


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230 


LABORATORY METHODS OF STUDYING GUN EROSION 


tens of molecules thick. Blocks of gun steel held in 
recesses in the side walls of the explosion vessel, for 
example, acquired such thin films of reaction prod¬ 
ucts that even fair diffraction patterns were extremely 
difficult to obtain. 

The results of the studies on blocks of gun steel are 
given in Section 12.2.2. 

112 5 Circular Vents in Stressed Blocks 

Stresses in guns might facilitate erosion by expos¬ 
ing fresh metal surfaces to the powder gases. Con¬ 
ceivably, this might result in continuously renewed 



Figure 9. Stressed erosion-vent apparatus. (This 
figure has appeared as Figure 1 of NDRC Report 
No. A-431.) 


chemical reaction or absorption of the powder gases. 
An experiment was set up to test such a possibility. 103 

Figure 9 shows a cross section of the apparatus in 
which a type of vent plug was used. The explosion 
vessel was cylindrical about the vertical axis. Each 
section shown, except that of the specimen, is cylin¬ 
drical about either the vertical or the horizontal axis 
of the containing chamber. All parts* were made of 
steel except the brass rupture disk R and the copper 
and rubber obturating disks A and A- Pressure was 
applied from a hydraulic press to the parts in the ver¬ 
tical chamber and was transmitted to the ends of the 
specimens through the hardened steel cylinders A 


and C 2 , the copper and rubber obturating disks A 
and A, and two additional hardened cylinders C 3 and 
C 4 . The specimen was rectangular in cross section, as 
shown in Section A-A, and the space between it and 
the cylindrical chamber wall was filled by two steel 
blocks Bi and B 2 which occupied almost the entire 
space but which did not interfere with the compres¬ 
sion of the specimen. With this arrangement the cyl¬ 
indrical surface of the hole was in radial compression 
along the line of intersection with a vertical diametral 
plane and in radial tension along the intersection with 
a similar horizontal plane. 

The horizontal chamber defined the path of the gas 
stream. The charge was contained in a caliber .30 
cartridge case which was inserted into the powder 
chamber P of a caliber .30 Springfield breech mech¬ 
anism attached to one end of the horizontal chamber. 
A small replaceable plug X narrowed the gas stream 
and led it to the entrance of the vent V. After passage 
through the vent, the gases entered a small chamber 
Y which was terminated by the rupture disk R. The 
gas pressure built up behind this disk until it broke, 
permitting the main flow of gas to pass through the 
vent. 

The ballistic conditions chosen for the tests were 
obtained with a vent of % in. diameter, a charge of 
2.5 g of powder NH-M1 (for 37-mm gun M1916), a 
loading density of 0.47 g/cc, and a brass disk that 
ruptured under a static pressure of 2,000 atm in cal¬ 
ibration tests. The applied stresses on the cross sec¬ 
tion of the specimen corresponded to pressures in the 
range of 16,600 to 83,100 psi. 

After firing, the diameter of the vent in different 
directions was measured with a traveling microscope. 
Any differential change in the diameter of the vent in 
the vertical or horizontal position indicates a differ¬ 
ent degree of erosion under tension and under com¬ 
pression, respectively. 

As further discussed in Section 13.4.1, the measure¬ 
ments showed that there was no appreciable differ¬ 
ence in the diameter for surfaces under tension or 
under compression. 

112 6 Collection of Particles and Gases 

The collection of solid particles and gases dis¬ 
charged from a small-caliber weapon was facilitated 
with the design of the collection tube 20 shown in Figure 
10. This tube, 4 ft long, was made of Pyrex glass with 
a 0.12-in. wall and 4-in. diameter. It was hermetically 
sealed at the two ends by heavy brass caps. The muz- 


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OTHER LABORATORY EROSION EXPERIMENTS 


231 



Figure 10. Glass tube for collecting gaseous and solid material ejected from the muzzle of a caliber .30 rifle behind 
the bullet. (This figure has appeared as Figure 1 of NDRC Report No. A-93M.) 


zle of a caliber .30 rifle barrel, held in a Mann rest, 
could be inserted in a gas-tight manner into one cap. 
When the rifle was fired, the bullet passed down the 
collection tube and emerged through a renewable 
disk of thin metal covering a central hole in the cap at 
the target end of the glass tube. A removable optical 
system 20 in the bore of the rifle afforded a means of 
sighting the line of fire with the requisite accuracy. 

Exit and entry ports were provided by a radial hole 
in each of the brass caps. One of these ports was used 
for evacuating the system. A little heavy grease on 
the neck of the cartridge case completed the vacuum 
sealing of the system. The pressure in the collecting 
tube before firing was reduced to at least 1 cm of Hg, 
otherwise the mortality rate of the glass tubes was. 
high. In special experiments the pressure was reduced 
to 1 mm of Hg or less. 

The other port could be used to flush out the sys¬ 
tem with a gas, such as nitrogen, or to permit the 
collected gases to expand into an evacuated bottle. 
The latter procedure provided extra room for the 
gases from high-pressure charges and was used at 
times in order to collect gas samples. In collecting gas 
samples the hole left by the bullet was closed imme¬ 
diately by a wad of Plasticene applied by the oper¬ 
ator. This procedure was crude, but the apparatus 
was not designed or used for the refined collection of 
powder gases. 

The chief uses of this apparatus are described in 
Sections 14.5.3 and 16.3.2. Filings of various metals 
or nitrides were mixed with the powder charges, 
otherwise the rounds were standard in all respects. 
The filings were recovered from the collecting tube, 
after firing, and examined by x-ray analysis for alter¬ 
ation or formation of reaction products. 28 - 79 

Some subsidiary experiments were performed with 


this apparatus in connection with the use of tracers 
for sulfur 53 and nitrogen. 70 These experiments were 
devised to study the distribution of the tracer ele¬ 
ments in the powder gases, as described in Sections 
14.4 and 14.5, respectively. 

113 OTHER LABORATORY EROSION 
EXPERIMENTS 

11,3,1 Vent Plugs Subjected to 

Adiabatically Compressed Gases 

If a gas at room temperature and 1 atm pressure is 
compressed adiabatically to about 1,000 atm its tem¬ 
perature will be increased by 1000 C to 3000 C. The 
actual rise of temperature will depend on the nature 
of the gas. 

This phenomenon opens up the possibility of study¬ 
ing the erosive effect of the individual components of 
powder gas and also of differentiating between chem¬ 
ical and purely thermal effects by using an inert gas 
such as argon. With proper design of apparatus it 
should be possible to control within certain limits the 
important parameters of time, temperature, and pres¬ 
sure so that they will be of the same order of magni¬ 
tude as those in a gun. 

An apparatus 101 that was used in some preliminary 
experiments of this nature is illustrated in Figure 11, 
which shows the assembled compressor and vent 
plug. The gases were compressed by the piston, which 
was driven by a falling weight. They were expelled at 
the bottom of the cylinder through a hole in the side. 
The bottom of the cylinder was sealed with a piezo¬ 
electric gauge protected by a steel disk (not shown). 
A slot in this disk was lined up with the escape hole 
so that the piston would not block the efflux of the 


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232 


LABORATORY METHODS OF STUDYING GUN EROSION 


gases. The escaping gases passed through a vent plug 
that consisted of a small hole in a washer of the de¬ 
sired material. In some experiments where flat sur¬ 
faces were desirable, a slit-type vent was formed by 
mounting two small rectangular blocks of the mate¬ 
rial to be tested. 

The experiments carried out with this apparatus, 
which are described in Section 14.6, showed its lim¬ 
itations. They also led to a conclusion concerning the 
desirable form of a more powerful apparatus, which 
is illustrated in Figure 16 of Chapter 14. 



Figure 11. Preliminary form of apparatus for adi¬ 
abatic compression of gases (see Figure 16 in Chapter 
14 for proposed improved form). (This figure has ap¬ 
peared as Figure 2 of NDRC Report No. A-429.) 

113 2 Electron Bombardment Apparatus 

An experiment was devised for the purpose of eval¬ 
uating the effects of heat alone on the bore surface. 104 
Electrons from a hot tungsten filament were made to 
strike a specimen of gun steel or other material caus¬ 
ing the surface of the specimen to be heated to a high 
temperature for a very short time. The use of electron 
bombardment proved to be amenable to the desired 
degree of control of time and intensity of heating. 

The specimen, mounted on a water-cooled column 
as shown in Figure 12, was a rod % in diameter 
with the end to be exposed carefully polished. This 
rod formed the anode of a gaseous triode contained in 
a bell jar under a vacuum of 5 X 10 -3 mm of mercury, 
as shown diagrammatically in Figure 13. The use of 


nitrogen or argon at these pressures served to reduce 
the space charge. The specimen was insulated with a 
Pyrex shield so that only the polished end was effec¬ 
tive as the anode. The circuit constants were chosen 
so that the time of exposure to the bombardment 
would be about 0.01 sec, which is comparable to the 
time of exposure of the bore surface to powder gases 
at the origin of rifling of a 3-in. gun. Electron dis¬ 
charges, usually about 1,000 in number, each lasting 
about 0.01 sec, were repeated at 3-sec intervals. Power 




Figure 12. Metallic sample mounted as an anode for 
subjecting it to thermal shock by means of electron 
bombardment. (See Figure 13 for diagram of electrical 
circuit of electron bombardment apparatus.) (This fig¬ 
ure has appeared as Figure 5 of NDRC Report No. 
A-432.) 


amounting to 20 kw cm -2 was applied with each pulse. 
It was calculated that this amount of energy would 
be required to raise the surface temperature to the 
melting point of steel in 0.01 sec. 

The results obtained in the course of the develop¬ 
ment of the apparatus have been reported. 104 It was 
hoped originally that the bombardment by electrons 
would provide a simple method for evaluating mate¬ 
rials in terms of the purely thermal effect as mentioned 
in Section 16.3.1. However, an erratic behavior of the 
discharges, despite very careful control of gas pres¬ 
sure, filament current, and anode voltage, resulted in 
insufficient accuracy and too poor a degree of repro- 


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OTHER LABORATORY EROSION EXPERIMENTS 


233 



Figure 13. Diagram of electrical circuit of electron bombardment apparatus. (See Figure 12 for details of mounting 
of sample.) (This figure has appeared as Figure 4 of NDRC Report No. A-432.) 


ducibility for quantitative evaluation in the hoped- 
for manner. Better control of the discharge might be 
obtained with a high-vacuum tube instead of the 
gaseous triode used. 

As a guide to continuation of this work in the 
future it is well to list the principal difficulties en¬ 
countered. It was difficult to insulate the specimen, 
and therefore arcing between the specimen and insu¬ 
lator was a frequently objectionable phenomenon. 
This probably arose from an accumulation of the 
electrons which struck the specimen. The discharge 
did not act uniformly on the surface but tended to 
concentrate on some areas of the specimen, leaving 
other parts relatively unaffected. This irregularity 
vitiated to some extent the calculation of the peak 
temperature attained by the surface. 

Approximately 300 specimens were exposed to elec¬ 
tron bombardment. Photographs of the surface and 
microphotographs of etched cross sections were made. 
Although the investigation did not result in the de¬ 
velopment of a simple quantitative method of evalu¬ 
ating materials with respect to the action of heat 
alone, one result was evident — that, although crack¬ 
ing was observed on specimens fired in nitrogen, none 
of the specimens exposed showed cracking of the kind 
observed in guns. This result seems to indicate that 


the cracking observed on a gun bore surface is not 
caused by heat alone. This important conclusion is 
discussed in Section 13.5.3. 

11,3,3 Apparatus for Cavitation Erosion 

Cavitation erosion was one of the phenomena stud¬ 
ied in an attempt to evaluate the erosion resistance of 
different materials, as described in Section 16.3.3. 
Apparatus already available at the Armour Research 
Foundation was used for the purpose. 38 The specimen 
was submerged in water in the lower end of a nickel 
tube. This tube was set in vibration by magnetostric¬ 
tion at a frequency of 8,000 c and an amplitude of 
about 0.05 mm. After a definite length, of time, the 
specimen was dismounted and examined with a bin¬ 
ocular microscope to determine the extent of erosion. 
On the basis of visual examination, the metals were 
listed according to the degree of resistance to cavita¬ 
tion erosion. 

The results showed that the materials that were 
most resistant to vent-plug erosion, were least resis¬ 
tant to cavitation erosion, and vice versa. The ob¬ 
vious conclusion, therefore, was that vent or gas ero¬ 
sion and cavitation erosion are not related, and that 
the apparatus is not useful to determine the resistance 


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234 


LABORATORY METHODS OF STUDYING GUN EROSION 


of metals and alloys to gas erosion. However, inas¬ 
much as muzzle erosion seems to be the result of 
mechanical wear, a better correlation might be ob¬ 
tained between cavitation erosion and muzzle erosion. 
Whether such a correlation exists has not been deter¬ 
mined. 

114 DISENGAGEMENT AND SEGREGATION 
OF EROSION PRODUCTS e 

11,4,1 Introduction 

In order to make chemical analyses of the products 
of erosion found on the bore surfaces of guns, it is 
necessary to disengage them from the unaltered steel. 
Analysis of the segregated material by a combination 
of chemical and x-ray techniques gives far more infor¬ 
mation concerning the nature of the alteration of bore 
surfaces than does x-ray examination of the products 
in situ, since the latter techniques can only be used to 
determine what crystalline species are present. More¬ 
over, identification by the x-ray diffraction method of 
the compounds entrapped in the coppering of guns is 
not possible unless they are removed from the 
copper. 

In one procedure, described in Section 11.4.2, the 
bore surface products were dislodged mechanically. 
In the other two, chemical methods were employed. 
In one of them the erosion products entrapped in the 
copper were dislodged by preferentially dissolving the 
latter, as described in Section 11.4.3. In the other, the 
erosion products covering the steel surface were dis¬ 
lodged by using a reagent that dissolves the under¬ 
lying steel. This is discussed in Section 11.4.4. 

Preliminary to determining the carbon and nitro¬ 
gen content of reaction products formed in a gun bore, 
it was of course necessary to remove grease and par¬ 
ticles of unburned powder from the specimens. Ordi¬ 
narily this was done simply by washing them with pe¬ 
troleum ether to remove grease and with acetone to 
remove powder particles. The grease was often ten- 
aceously lodged in the fine crack system and was 
therefore very difficult to remove. In order to insure 
complete removal it was sometimes necessary to use 
a Soxhlet extractor. The specimen was removed after 
each extraction and sufficient time was given for the 
solvent to seep out of the cracks by capillary action 

e This section has been condensed from sections of NDRC 
Report A-426, 98 in which complete details for carrying out the 
different procedures are given. 


and thus bring the grease to the surface where it 
could be readily removed by more solvent. 

11.4.2 Mechanical Removal of 

Erosion Products 

The following procedure, which serves as an ex¬ 
ample of removal of products by mechanical means, 
was used on two quarter portions of a section of a 
3-in. gun liner 98 extending 5.5 to 9 in. from the origin 
of rifling. The pieces were mounted on a mandrel in a 
lathe and gun steel was removed from the outside 
surface until a thickness of 2 mm was obtained. These 
thin quarter sections were then placed bore surface 
downward and were flattened out by hammering. The 
progressive stress-damage cracks (defined in Section 
13.5.3) greatly facilitated the operation and, in fact, 
made it possible to separate the lands and grooves. 
The specimens were then thoroughly washed with 
petroleum ether to remove grease and with acetone 
to remove powder particles, as described above. 

Solid reaction products were removed mechanically 
from areas that formed the walls of these major 
cracks and were analyzed chemically (for carbon and 
nitrogen), spectrographically, and by x rays. In addi¬ 
tion to removing the crack-filling material, it was pos¬ 
sible to dislodge mechanically metallic “beads” that 
projected over the edges of the cracks. These “beads” 
were examined by x rays. The results of the analysis 
of the crack-filling material and of the “beads” are 
given in Sections 12.4.1 and 12.4.2, respectively. 

11.4.3 Removal of Copper and 
Entrapped Erosion Products 

The “copper” that occurs in the central portion of 
a gun bore, as described in Sections 10.5.4 and 12.3.1, 
provides a “storehouse” for products of erosion that 
have been carried forward. A method was developed 
for removing this copper or gilding metal without at¬ 
tacking the underlying steel or the erosion products. 
Such a technique not only affords a means of separat¬ 
ing the erosion products but also permits examination 
of the underlying bore surface. 

Procedure 

Specimens to be decoppered were treated with an 
ammoniacal solution of ammonium carbonate and hy¬ 
drogen peroxide at about 5 C. The relative amounts 
that were found to be expedient were 15 ml ammonia 


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DISENGAGEMENT AND SEGREGATION OF EROSION PRODUCTS 


235 


water (28-29% NH 3 ), 0.5 g of powdered ammonium 
carbonate, and 3 ml of 30% hydrogen peroxide. Such 
a solution is sufficient to dissolve about 0.4 g of 
copper. It is important that no water be added to the 
decoppering solution; otherwise attack of the steel 
takes place. It is also important that the temperature 
of the solution and steel be held between 0 and 5 C; 
for if it is inadvertently allowed to warm to room 
temperature, oxidation of the iron takes place, which 
is revealed by clouding owing to a precipitation of a 
hydrated oxide of iron. 

Neither steel nor chromium plate was appreciably 
attacked in a number of experiments as long as the 
above precautions were observed. Polished specimens 
were still bright after several hours immersion in the 
ice-cold solution. Very few of the reaction products 
formed by the interaction of the powder gases and 
steel seem to be attacked within the time that they 
are usually exposed to the solution. Even ferrous ox¬ 
ide, present as wustite, is not decomposed. 

A thin coating of a few microns of copper on a rela¬ 
tively smooth surface was removed in 2 or 3 min, but 
a coat which is one or more millimeters thick may 
take many hours and require several treatments with 
fresh solution. If the surface is deeply eroded, the 
copper that is lodged in the bottom of the cracks ex¬ 
poses a much smaller surface to the solution than 
does copper spread out over the bore surface and may 
require repeated treatment for complete removal. 

Treatment of Residues 

To recover the gun-erosion products entrapped in 
the coppering, the mixture was transferred to a con¬ 
tainer surrounded by ice. The specimen was washed 
with concentrated ammonia water and treated again, 
if necessary, with the decoppering solution. The solid 
particles were promptly separated from the spent so¬ 
lution by centrifuging and were washed in the tube 
with chilled concentrated ammonia, and finally with 
“C.P.” alcohol. Four washings with each reagent 
were found to be sufficient. The bulk of the alcohol 
after centrifuging was decanted as completely as pos¬ 
sible, and the remainder removed by cautiously heat¬ 
ing at about 55 C, whereupon the tube was placed in 
a vacuum desiccator. The thoroughly dried residue 
was then ready for x-ray, microscopic, and chemical 
analysis. If this procedure is followed, the cementite 
frequently present in the residues is not attacked, 
but if the residue is allowed to remain overnight 
in contact with the decoppering solution at room 


temperature, it oxidizes and little or no cementite 
remains. 

WpJI 

Analysis of Solutions 

The solutions that have been separated from the 
residue may be evaporated to dryness in a porcelain 
or platinum basin and any lead, zinc, or sulfate that 
may be present separated from the copper. Analysis 
of the solution quickly yields information as to 
whether the projectiles fired from the gun being ex¬ 
amined had had copper or gilding metal rotating 
bands. 

114 4 Disengagement of Erosion Products 
by Attacking the Steel 

Another type of segregation of erosion products 
from the bore surface involved the use of a copper 
potassium chloride solution. This is a solution which 
dissolves the ferrite from beneath the erosion prod¬ 
ucts, which may then be analyzed and examined by 
a combination of techniques. In most instances, as 
discussed in Section 12.5.2, the products consisted 
mainly of cementite (Fe 3 C). Carbides of alloying ele¬ 
ments and nitrides were also found. 

Procedure 

The reagent was prepared by dissolving 300 g pure 
copper potassium chloride (CuCl 2 -2KCb2H 2 0) in 
a mixture of 65 ml of concentrated hydrochloric acid 
and about 800 ml of water. The solution was filtered 
through asbestos and diluted with water to a volume 
of one liter. 

The reagent was used at room temperature and 
was mechanically stirred to accelerate decomposition 
of the steel. During this decomposition, copper was 
deposited on the specimen but was redissolved by the 
time the ferrite was completely dissolved. Cuprous 
chloride was also formed; it is soluble in the reagent 
but somewhat insoluble in water. Washing of all speci¬ 
mens and segregated residues, therefore, was done 
first with portions of the fresh solution before washing 
with water. 

The reagent was applied directly to the bore sur¬ 
face of a gun section, remaining in contact only long 
enough to dissolve the immediately underlying steel. 
The flakes were then picked off. Since the steel was 
not completely decomposed in this case, the copper 
deposited in the course of the reaction was not dis- 


CONFIDENTIAL 



236 


LABORATORY METHODS OF STUDYING GUN EROSION 


solved and had to be removed by the decoppering 
technique previously described in Section 11.4.3. 

Relative Insolubility of Compounds 
in the Reagent 

The success that was obtained with such a reagent 
in segregating cementite and other bore-surface prod¬ 
ucts was due to the fact that they have greater re¬ 
sistance to attack by the reagent than has iron. Thus 
the erosion products are not “insoluble’’ in the re¬ 
agent in the sense that the term “insolubility” is 
normally used. It is important that the reagent not 
be allowed to remain in contact with the specimen 
for too long a time as it dissolves erosion products 
other than cementite. Eventually even cementite 
may decompose in the solution. 

Products other than cementite were segregated 
from steel and from eroded bore surfaces by means of 
this reagent. Thus from a plain carbon steel no ce¬ 
mentite was obtained but instead alabandite, a crys¬ 
talline form of manganese sulfide, together with a 
highly carbonaceous residue. Nitrides were frequent¬ 
ly segregated from the eroded bore surfaces. Their 
relative insolubility in the reagent is discussed in Sec¬ 
tion 12.5.2. 

Even more erosion products were segregated with 
a solution of copper potassium chloride to which no 
acid had been added. However, quantitative analysis 
of them was made impossible owing to rapid hydroly¬ 
sis of iron with formation of a presumably hydrated 
oxide of iron. The non-acid solution, however, was 
useful when only x-ray analysis of the products was 
required. 

Novel Method of Studying Crack Systems 

The crystals of cementite in outer white layers that 
had been removed by means of the copper potassium 
chloride solution (see Section 12.5.1) were so coherent 
that when care was exercised they could be removed 
as scales which comprised a cast of the reticulated 
crack system. A refined method of obtaining such a 
cast consisted of coating with a Vinylite resin the 
eroded surface of a thin section taken parallel to the 
bore surface and then dissolving the ferrite away 
from the other side, whereupon the flakes adhered to 
the Vinylite. The Vinylite coating and the adherent 
erosion products was called a “plaque.” Stereoscopic 
photographs of a number of plaques were taken. One 
of a pair of photographs is reproduced in Figure 14. 


Plaques may prove useful in studies of crack systems. 
They should not be confused with the replica films 365 
that have been used to study the eroded surface of a 
gun bore. 

n.4.5 Fractionation of Erosion Products 

Erosion products disengaged from the bore surface 
were sometimes separated from one another by elu- 
triation in an organic solvent, such as alcohol. Two 
refinements of the ordinary elutriation process were 
employed. The particles stirred up in the liquid were 
separated (1) by centrifuging the suspensions for dif¬ 
ferent lengths of time, and (2) by holding a magnet 
near the bottom of the container when decanting the 
liquid. 

115 TECHNIQUES USED TO EXAMINE 
THE PRODUCTS OF EROSION 

1151 Visual and Metallographic 
Examination of Bore Surfaces 

The visual examination of eroded bore surfaces is 
not a technique in the same sense as is examination 
by x-ray or electron diffraction. Much depends on the 
acuteness of observation of the investigator, his 
knowledge of what might be expected to happen in 
guns, and his general experience in carefully examin¬ 
ing gun bores. In short, he must be a detective. His 
results are not only difficult to describe but also they 
are sometimes difficult to photograph in such a way 
as to present to the reader what the investigator him¬ 
self sees. 

The Greenough-type stereoscopic microscope has 
been used 112 recently to relate the topography of the 
bore surface with positions in the gun, with various 
structures and constituents of the surface, and with 
various processes and stages of erosion. The typical 
surface features of gun bores are more readily inter¬ 
preted when observed stereoscopically than when 
seen in photographs, such as those shown in Section 
10.5.2. 

Metallographic examination of sections cut normal 
to the bore surface when polished and etched reveals 
the character of the surface, particularly the layers 
which are present. Different etchants have been used 
to show the constituents present in these layers which 
are described in Section 12.1.2. 

Careful microscopic observation of the altered 
layer formed in the caliber .50 erosion testing gun 


CONFIDENTIAL 



TECHNIQUES FOR EXAMINING PRODUCTS OF EROSION 


237 


have led to interesting conclusions concerning its role 
in gun erosion. 124 The salient features of these con¬ 
clusions are given in Section 13.2.3. 

1152 Examination by X-Ray and 
Electron Diffraction 

Relative Advantages of the Two Techniques 

Examination of eroded gun bores and test spec¬ 
imens by means of x-ray and electron diffraction has 
yielded much information concerning chemical sub¬ 
stances formed during erosion. The special utility of 
these methods is due to the fact that only a minute 
quantity of material is necessary for study. 

By the use of electron diffraction notable success 
has been attained in the identification of films of 
crystalline material which are too thin to be ex¬ 
amined by x-ray diffraction. It was possible to iden¬ 
tify successive altered layers on steel blocks exposed 
to powder gases in the apparatus described in Section 
11.2.4 by combining this technique with one devel¬ 
oped for the removal of exceedingly thin films. 31 

X-ray examination yields more information when 
thick films are being examined, as x rays have 
greater penetrating power than do electrons. X-ray 
diffraction has been applied more extensively to stud¬ 
ies of erosion than has electron diffraction, because a 
smooth surface is required for the latter method. The 
cracked bore surfaces of eroded guns are usually too 
irregular for examination by-electron diffraction. 

Electron Diffraction Technique 31 > 137 

In examination by electron diffraction, the eroded 
specimen is placed in an exhausted container and 
electrons of uniform energy are caused to impinge on 
the surface so that they are diffracted from it at a 
very small glancing angle, about 0.5 degree. The re¬ 
sulting ringlike diffraction pattern is then recorded on 
a photographic plate. The identity of one or more 
substances present is established by comparing the 
measured diameters of the rings with those of stand¬ 
ard specimens. Only rough estimates of the relative 
amounts of the constituents may be obtained from 
the diffraction pattern. 

In order to obtain a good diffraction pattern, the 
surface must not be microscopically smooth; there 
must be projections on the surface of small enough 
dimensions for the electrons to go clear through. 
Otherwise, refraction of the electrons at the small 


angle of incidence prevents formation of the diffrac¬ 
tion pattern. 

Some eroded specimens do not meet this require¬ 
ment, and it is therefore necessary to resort to some 
method of roughening the surface in order to obtain 
a satisfactory diffraction pattern. If the chemically 
altered layer is thick enough so that there is no 
danger of removing too much of it, the surface may 
be scratched with emery. For thin layers, however, 
scratching with emery paper, brushing with emery 
powder, or blowing with the powder is unsatisfactory. 



A B C 


Figure 14. Plaque of an area of eroded 3-in. gun liner 
No. 1460 near the origin of rifling. Groove extends from 
A to B, and land from B to C; 2X. (This figure has 
appeared as Figure 1A of NDRC Report No. A-426.) 

To overcome this difficulty a ruling machine was 
built, using a special design of sapphire phonograph 
cutter so mounted that it digs up the surface in much 
the same way that a plow digs up the earth. This 
“microplow” (Figure 15), rules 8,000 “furrows” per 
inch, as shown in Figure 16, the furrows being of the 
order of 1 n in depth. 

Thin flakes, removed in etching altered surfaces, 
were examined by allowing the electrons to pass 
through them. 

Another limitation of the electron-diffraction tech¬ 
nique is that the surface on which the electron beam 


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238 


LABORATORY METHODS OF STUDYING GUN EROSION 


is directed must be flat or nearly so. Thus, in the ex¬ 
amination of a section of a ring of the bore surface, 
the beam cannot be in the plane of the curved speci¬ 
men. This difficulty diminishes, of course, for guns 
of larger caliber. 

In studies of gun erosion, electron-diffraction ex¬ 
amination has been applied to gun specimens, blocks 
exposed to carbon monoxide-oxygen mixtures (Sec¬ 
tion 14.3), blocks heated by adiabatically compressed 
gases (Sections 11.3.1 and 14.6.6), specimens used in 
the study of the effect of sulfur and other components 
of black powder (Section 14.4.4), and specimens 
bombarded with electrons (Section 11.3.2). 

X-Ray Diffraction Technique 

Examination by x-ray diffraction is a technique 
similar to that of electron-diffraction analysis, which 


was described above. Surface layers may be examined 
in situ by the glancing angle method, using an angle 
of 10 to 30 degrees. Powder photographs may be 
made of loosely held or dislodged materials. 

In some cases where great difficulty in obtaining 
clear patterns is experienced, it has been recom¬ 
mended that x rays of various wavelengths be used. 137 
Variation of the wavelength is obtained by using 
x-ray tubes having targets of different metals. Each 
of the elements of the periodic table has a character¬ 
istic K absorption wavelength for x rays. If the x rays 
used to examine a specimen have a wavelength shorter 
than its K absorption limit, the pattern has a heavy 
background caused by fluorescence, and the lines of 
the pattern are weak due to the heavy absorption. 
If the x rays are just a little longer than the K ab¬ 
sorption limit, the resulting pattern has much less 
background and the pattern stands out clearly. 



Figure 15. Microplow—an apparatus for ruling a series of fine lines on a specimen of eroded gun steel preparatory to 
examination by means of electron diffraction. (This figure has appeared as Figure 1 of NDRC Report No. A-465.) 


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TECHNIQUES FOR EXAMINING PRODUCTS OF EROSION 


239 


As an example of this method of absorption anal¬ 
ysis, suppose the specimen is rich in iron, which has 
a K absorption limit at 1.739 A. If x rays of wave¬ 
length 1.539 A from a copper target are used, the 
resulting pattern is very poor with a heavy back¬ 
ground. If, however, radiation of wavelength 1.934 A 
from an iron target is used, the pattern is stronger 
with much less background. If it is not known that the 
specimen is rich in iron, this result only indicates 
that the specimen is rich in some element with an 
absorption limit greater than 1.539 A but less than 
1.934 A. This restricts the choice to iron, cobalt, and 
manganese. 

The situation is not so clear cut if a pattern ob¬ 
tained with x rays from the copper target is as good 
as or better than the one taken with iron radiation. 
About all that can be said in this case is that the 
specimen is not rich in iron, cobalt, or manganese; 
and if the pattern taken with iron radiation is a good 
clear one without much background, then the speci¬ 
men is probably also not rich in chromium, vana¬ 
dium, titanium, or scandium. 

Unidentified Lines. It will be seen in Chapter 12 
and elsewhere 28,5 °’ 109 ’ 123 ■ 137 - 261 that reference is fre¬ 
quently made to the presence of unidentified lines 
in the photographs obtained by x-ray and electron 
diffraction. Unfortunately, there are insufficient data 
at hand to permit one to assign these lines to definite 
compounds. Then too, the effect of the presence of 
impurities in causing distortion of lattices also re¬ 
quires further study. When such data have been ob¬ 
tained, x-ray examination will become an even more 
powerful tool for the examination of eroded speci¬ 
mens, with the result that an even better understand¬ 
ing of the chemical aspects of gun erosion is likely to 
be obtained than is at present possible. 

115 3 Chemical Analysis 98 

Introduction 

In addition to and in conjunction with analysis of 
bore-surface products by x-ray and electron diffrac¬ 
tion, chemical analysis, both qualitative and quan¬ 
titative, has been employed. While x-ray and elec¬ 
tron diffraction identify particular crystalline phases, 
chemical analysis tells only what elements and how 
much of them are present. It is useful however, in 
showing up elements that give rise to electron and 
x-ray diffraction patterns that cannot be identified 
or that are constituents of compounds whose pres¬ 
ence cannot be detected by diffraction techniques 


because they are either amorphous or not present in 
sufficient quantity. Moreover, analysis by diffraction 
methods tells nothing about the purity of the com¬ 
pounds identified. For example, in the study of the 
carbide segregated from the steel of a 3-in. gun liner, 
x-ray analysis showed that it had the structure of 
cementite (Fe 3 C). Quantitative chemical analysis, 
however, showed that the carbide from the steel con¬ 
tained appreciable amounts of the alloying constitu¬ 
ents, as shown in Table 2 in Chapter 12. Chemical 
analysis alone only rarely yields sufficient informa¬ 
tion to permit one to say whether a single substance 
or a mixture is present. 



Figure 16. Photomicrograph of a ruling made with 
the microplow shown in Figure 15; 350X. (This figure 
has appeared in Figure 2 of NDRC Report No. A-465.) 

Methods 

For quantitative analyses of erosion products, 
standard methods were employed. When the amounts 
of available material were small, as they usually were, 
these methods were modified to bring them within 
the bounds of semimicro procedures. It was desir¬ 
able to analyze all segregated products for carbon 
and nitrogen 49 in order to obtain information on the 
relative importance of these particular elements in 
the powder gases in relation to the causes of gun ero¬ 
sion. Very small amounts of iron were determined by 
potentiometric titration 21 of the electropolishing so¬ 
lutions used in connection with the carbon-penetra- 


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240 


LABORATORY METHODS OF STUDYING GUN EROSION 


tion studies described in Section 14.2. This procedure 
was also used for some of the segregated residues. The 
usual colorimetric methods were used to determine 
in erosion products the amounts of some of the ele¬ 
ments which are used as alloying constituents in gun 
steel. In order to determine exceedingly small 
amounts of molybdenum, a considerable refinement 
of the method 98 was necessary. 

The possible role of sulfur in gun erosion, discussed 
in Sections 12.3 and 14.4, led to the need for a means 
of determining this element in erosion products. The 
chemical analysis of a solution or residue can yield 
information as to the presence of sulfur but does not 
enable one to identify the compound of which it 
forms a part. In some cases, however, tests based on 
reactions in chemically sensitized gelatin-coated pa¬ 
pers can be employed to yield a fair amount of infor¬ 
mation with respect to the presence of sulfides. 

Such tests have long been used in metallurgical 
work to detect the presence of sulfides readily soluble 
in dilute sulfuric acid. A so-called Baumann print, 
which is made on silver bromide paper by wetting it 
with dilute sulfuric acid and then pressing it against 
the polished surface of the steel, reveals the acid- 
soluble sulfides by precipitation of silver sulfide in the 
gelatin coating of the paper. The general principle of 
the Baumann print may be extended. Thus silver 
bromide paper soaked in water may be used to de¬ 
tect the presence of water-soluble sulfides; and lead 
cyanide paper soaked in 5% potassium cyanide may 
be used to detect copper sulfides. 

All three of the above types of contact printing 
were employed in the study of erosion products. The 
use of these sensitized papers enabled the detection 
of sulfides which were not revealed by x-ray or elec¬ 
tron diffraction examination. 

The physical condition of the sulfide is a control¬ 
ling factor in obtaining good prints. Thus, precipi¬ 
tated zinc sulfide yields a good print, but the crystal¬ 
line form, sphalerite, gives no print. The same 
phenomenon has been observed with precipitated 
nickel and iron sulfides and their respective crystal¬ 
line forms, millerite and pyrrhotite. 

Details and applications of these printing experi¬ 
ments have been reported. 53,95 

11,5 4 ^ Method for Determination of the 
Melting Temperatures of Gun 
Erosion Products 

One of the methods of determining bore-surface 


temperatures described in Section 5.4.3 was an ex¬ 
perimental method 105 based on observations that 
some of the products of erosion found on gun bore 
surfaces showed features that indicated they had 
been in a molten state at some stage during firing. 
This method was used to determine the incipient 
melting points of some of the erosion products that 
were disengaged and segregated by the techniques 
described in Section 114. 

11,5,5 Bore Surface Reactions Studied 
with Isotopic Tracers 

The recent methods of concentrating the less abun¬ 
dant isotopes of the elements and of producing radio¬ 
active isotopes have provided new and powerful 
means of tracing individual elements, in particular, 
nitrogen, carbon, and sulfur, in chemical and physical 
processes. In some of the experiments on bore sur¬ 
face reactions the ordinary chemical and physical 
methods would not have provided a solution to several 
pertinent problems. These problems were successfully 
studied by the very sensitive method of adding a 
tracer isotope, either one of the less abundant stable 
isotopes or a radioactive isotope, as described in 
Chapter 14. 

The work with tracers is largely complementary to 
the study of eroded bore surfaces described in this 
chapter because it demonstrated that the process of 
chemical deterioration of the bore surface is a contin¬ 
uous one commencing with the first firings in a new 
barrel. 

The use of tracers to study the processes of diffu¬ 
sion and reaction of the constituents of the powder 
gases below fresh bore surfaces of steel is more illu¬ 
minating than an attempt to demonstrate the course 
of such processes in badly eroded surfaces. In many 
of these experiments with new bore surfaces the con¬ 
stituent was present in too small an amount for iden¬ 
tification or separation by ordinary chemical proce¬ 
dures. 

Demonstration of Equilibrium. For the study of 
erosion the tracer isotope was introduced into the 
powder charge for a gun by incorporating it as a coat¬ 
ing in the form of a compound which would be com¬ 
pletely dissociated when the powder burned. Sub¬ 
sidiary experiments were performed to demonstrate 
a uniform distribution of a tracer element among the 
appropriate constituents in the powder gas, a neces¬ 
sary step before interpreting measurements on the 
bore surface. From one such experiment came a 


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TECHNIQUES FOR EXAMINING PRODUCTS OF EROSION 


241 



Figure 17. A section of a caliber .30 barrel mounted as a Geiger-Mueller counter for measuring the beta ray activity 
of a radioactive tracer-element incorporated in the bore surface. One half of the section has been cut away to show 
the Lucite fittings. (This figure appeared as Figure 1 of NDRC Report No. A-427.) 


method 61 of determining the state of equilibrium 
among the carbon atoms of a propellant gas. This 
is discussed in Section 2.3.5. 

Sampling. After the gun had been fired, the tracer 
atoms that had reacted with the bore surface needed 
to be recovered so that the proportion of them could 
be determined. In the study of the penetration of 
nitrogen, which is described in Section 14.5, thin 
layers of steel were bored out of a caliber .30 barrel, 
dissolved in acid, and then analyzed by means of a 
mass spectrograph. In this way the relative abun¬ 
dance of the rare nitrogen isotope 15, with which the 
powder charge had been enriched, was determined. 

In the study of the reaction of carbon gases with 
the bore surface, described in Section 14.2, successive 
layers of the subsurface of the bore, which were often 
no thicker than }/% fx, were removed by a special elec¬ 
tropolishing technique in which the iron was pre¬ 
served for quantitative analysis. 22 The radioactivity 


of the surface was determined after each such re¬ 
moval, in order to determine the amount of the radio¬ 
active tracer, which was the long-lived radioactive 
isotope of carbon. The carbon content of each layer 
was thus indexed by the decrease in activity of the 
radiocarbon. The final result was the construction of 
a carbon-penetration curve on a microscale. Further 
details appear in Section 14.2.3. 

Isolation of the sample was simplified in one set of 
experiments 53 on the reaction of sulfur in the powder 
gases with steel by using steel test rods that were ex¬ 
posed to the powder gases in the apparatus described 
in Section 11.2.4. How much of the radiosulfur added 
to the powder charge remained on the surface of the 
test rod was determined by inserting the test rod 
directly into a specially designed Geiger-Mueller 
counter. The extent of the penetration, which is dis¬ 
cussed in Section 14.4, was determined by repeating 
the measurement of radioactivity after successive 


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242 


LABORATORY METHODS OF STUDYING GUN EROSION 


layers of surface products and steel had been removed 
on a metallographer’s finishing lap, the rods being 
weighed on a microbalance to determine the amount 
removed. 

Special Geiger-Mueller Counters. An inherent diffi¬ 
culty in many tracer experiments is the tremendous 
dilution of the tracer isotope. This was particularly 
the case in the experiments with radiocarbon just 
mentioned, for the amount of ordinary carbon was 
fixed by the size of a caliber .30 charge and only very 
little of the carbon gases penetrated or reacted with 
the bore wall. Furthermore, radiocarbon emits a very 
soft beta-ray and the counting of these electrons is 
very inefficient. 

Several of the experiments were possible only be¬ 
cause the efficiency of counting the soft beta-rays was 
maintained at a high level by employing a novel form 
of Geiger-Mueller counter. A half-section model is 
illustrated in Figure 17. The cathode of the counter 
was itself the specimen, namely a short length of the 
barrel from whose bore surface the beta-rays were 
emanating. This arrangement may be useful in other 


studies, as for example in examining the surface ef¬ 
fects of commercial carburization. 

In an early type of experiment in the study of the 
penetration and reaction of sulfur, a single round con¬ 
taining radiosulfur was fired from a caliber .30 rifle 
and the whole barrel was mounted as a Geiger- 
Mueller counter to measure the soft beta-rays from 
the radiosulfur. Such a counter worked, but is not to 
be recommended. Some information concerning the 
high reactivity of the sulfur gases with gun steel was 
obtained; but the exact location of the sulfur in the 
barrel was difficult to find and the amount of pen¬ 
etration could not be satisfactorily established. The 
procedure developed later for use with radiocarbon, 
described above, proved far superior. The only ad¬ 
vantage in using the whole barrel lies in the fact that 
its performance as a rifle is not impaired. The same 
barrel was used alternately in the rifle and as a Geiger 
counter in studying the effect of subsequently fired 
rounds. 

Temperature of Specimen. There would be little 
point to studying the penetration of carbon below the 



Figure 18. Assembly for firing a caliber .30 barrel at elevated initial temperatures. Note the apparatus for passing 
oxygen-free nitrogen through the barrel in order to protect the bore surface. The glass tube leading to the muzzle 
was removed about one second before firing. The gunner was protected by a sheet of I^-in. boiler plate, for the experiment 
was hazardous. Above all, smooth bolt action was imperative, because a “cook-off” would occur after a 2-second contact 
of the round with the hot chamber. The temperature controls and firing station were in the next room. (This figure has 
appeared as Figure 4 of NDRC Report No. A-427.) 


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TECHNIQUES FOR EXAMINING PRODUCTS OF EROSION 


243 


temperature of transition of alpha- to gamma-iron. 
Because the bore surface of the caliber .30 rifle barrel 
ordinarily does not reach this temperature, the test 
barrels were preheated in a protective stream of 
oxygen-free nitrogen. A photograph of the apparatus 
is shown in Figure 18. 

To gain some idea of the necessary amount of pre¬ 
heat, the question of simulating, in a small caliber 
barrel, the higher bore-surface temperatures and 
longer heating times involved in medium and large 
bore guns, was studied in some detail. Since it seems 
to be practicable to attain this purpose with respect 


to temperature and to some extent with respect to 
the significant times involved, this procedure may 
constitute a method of approach to other ballistic 
problems. 

Diagrams were prepared to compare the computed 
temperature-time curves for various guns and pro¬ 
pellants with similar curves for preheated caliber .30 
barrels." The penetration of carbon and the forma¬ 
tion of reaction products containing carbon was stud¬ 
ied in the abovamanner for a number of maximum 
bore-surface temperatures and propellants. The re¬ 
sults are summarized in Section 14.2.2. 


CONFIDENTIAL 



Chapter 12 

THE PRODUCTS OF GUN EROSION 

By E. G. Zies a and C. A. Marsh h 


121 SURFACE LAYERS IN ERODED GUNS 

i 2 .l i Investigations of Nature of 
Surface Layers 

E xamination of eroded guns has shown the ex¬ 
istence on the bore surface of layers that differ in 
several respects from the unaltered gun steel. It was 
soon recognized that in order to know what means 
could be employed to mitigate the deleterious effect 
of erosion, it would be necessary to learn what factors 
are involved in producing the altered bore surface, 
and how these factors are interrelated, as will be dis¬ 
cussed in Chapter 13. An essential part of such a 
study was a knowledge of the nature of the layers 
on the eroded bore surfaces. 

There are several distinct types of altered layers, 
not all of which are found in every gun. The layers 
found depend on a variety of factors. Not all of these 
variations were recognized by the early investigators; 
it is quite likely that they were not studying the same 
types of altered layers. 

The methods that have been used to study altered 
bore surfaces are numerous. They are of value only 
when the results are coordinated, so that one is sure of 
what type of altered layer is responsible for the ob¬ 
servations. The following techniques were employed 
by early investigators. 

Hardness tests showed that some of the bore sur¬ 
face layers were harder than the unaltered steel. The 
conclusions reached were that this hardness was due 
to one or more of the following processes: cold work¬ 
ing, martensite formed by quenching, introduction of 
carbon, introduction of nitrogen. The behavior of the 
altered layers on tempering led to the conclusion that 
there was no chemical alteration. Analyses of cuts 
taken parallel to the bore surface indicated in some 
instances that a carburized case had formed and in 
others that a nitrided case had formed. X-ray exam¬ 
inations of eroded bore surfaces revealed the presence 
of austenite but not of martensite or compounds of 

a Chemist, Geophysical Laboratory, Carnegie Institution of 
Washington. 

b Assistant Chemist, Geophysical Laboratory, Carnegie In¬ 
stitution of Washington. (Present address: U. S. Geological 
Survey, Washington, D. C.) 


iron. The retention of the austenite favored the 
theory of nitrogen penetration as a cause of erosion. 

Metallographic examination seemed to prove the 
same things that were found by the hardness tests, 
but etching seemed to show that no cementite was 
present. Most of the foregoing techniques have been 
markedly improved since the time they were first 
employed in the study of gun erosion. This early 
work has been evaluated elsewhere with respect to 
the limitations that prevailed at the time it was 
done. 16 ’ 261 

Kosting 261 initiated a truly systematic study of 
eroded guns. He was the first to show the complexity 
not only of the altered layers but also of the entire 
problem of gun erosion. Earlier investigators had 
used the term “white layer” to designate the whole al¬ 
tered zone, for with the usual, mild etching reagents 
employed in the metallographic examination of steels, 
this zone does not appear etched at low magnifica¬ 
tions. Kosting restricted the term “white layer” to the 
outer zone which constitutes only a minor part of the 
whole altered layer. In order to eliminate confusion, 
this practice is now generally followed. Moreover, the 
inner zone sometimes etches dark, thus rendering the 
term white layer a misnomer when applied to it. 


12,1,2 Description of Layers 

Introduction 

As was stated above, the layers that may be found 
in eroded guns vary. In some cases only the thermally 
altered layer that is common to all eroded guns is 
present. In the majority of large caliber guns, how¬ 
ever, all types of layers may be found if care is exer¬ 
cised to preserve the outermost layers in mounting 
and polishing the specimens. The following succession 
of layers is typical for guns of low-alloy steel that 
have been fired with single-base powders at normal 
pressures and rates of fire. The succession is from the 
bore surface inward to the unaltered steel: outer 
white layer or layers; inner white layer; thermally 
altered layer; troostite band. These features are 
shown in Figures 1 and 2. 


244 


CONFIDENTIAL 



SURFACE LAYERS IN ERODED GUNS 


245 


Thermally Altered Layer 

The thermally altered layer is common to all steel 
guns in which the bore surface has been heated above 
a critical temperature, about 720 C, for a critical 
length of time. 124 In some cases it is the only altered 
layer evident. In the usual case, however, it consti¬ 
tutes the bulk of the whole altered zone and over a 
considerable area of bore surface extends to a uniform 
depth below the surface. The thickness of this layer 
depends on many factors, among which may be men¬ 
tioned the caliber of the gun, the heat input to the 
bore, and the location in the bore. 


A B C D 



Figure 1. The typical succession of layers on the bore 
surface of an eroded gun: A—outer white layer; B—in¬ 
ner white layer; C—thermally altered layer; D—un¬ 
altered steel. [Cross section of 5-in./25-cal. gun tube 
No. 1982T at the origin of bore. Etched with picral 
( 500X )]. (This figure has appeared as Figure 15b in 
NDRC Report No. A-440.) 

The heat input, of course, depends on the flame 
temperature of the particular powder, the rate of fire, 
and the amount of powder. The correlation of the 
thickness of the thermally altered layer, heat input, 
and flame temperature from available data on tests 
with the caliber .50 erosion-testing gun is made in 
Section 15.5. 

It was also found in these studies carried out on the 
erosion-testing gun that the layer w r as thickest near 
the origin of rifling and that the thickness decreased 
muzzleward; in addition, the geometry of the lands 
caused the layer to be thicker under the land surface 
than under the groove surface. The depth of the 
thermally altered layer w f as practically independent 
of the number of rounds fired. In the case of chromi¬ 


um-plated barrels tested in this gun, thickness of the 
thermally altered layer depended on thickness of the 
chromium plate. As will be (^escribed in Section 31.5, 
experiments w^ere carried ]out to determine the 
minimum thickness of chromium plate necessary 
to suppress the fbrmation of a thermally altered 
layer. 

In some rapid-fire barrels the whole barrel except 
breech end is so hot and cools so slowly that a definite 
layer does not result. 

The exact nature of the thermally altered layer is 
still not known. So far no means of isolating it for 
study has been found. There is good reason for be¬ 
lieving that it has essentially the same chemical com- 



Figure 2. The typical succession of layers on the bore 
surface of an eroded gun. The outer white layer here is 
unusually thick, and the thermally altered layer has 
been partially tempered during firing. A—outer white 
layer; B—inner white layer; C—thermally altered 
layer; D—unaltered steel; E—troostite band. [Longi¬ 
tudinal cross section of 76-mm gun tube M1A1 No. 
1425 at the location of the mouth of the cartridge case. 
Etched with picral HC1 (1000X).] (This figure has ap¬ 
peared as Figure 13a in NDRC Report No. A-440.) 

position as the steel. For instance, when this layer is 
tempered, metallographic evidence shows it does not 
differ from the unaltered sorbitic gun steel. 261 Even a 
thin chromium plate, so long as it is intact, prevents 
the penetration of carbon and nitrogen into the steel. 
Nevertheless, a thermally altered layer can form un¬ 
der this thin plate. Studies of the penetration of car¬ 
bon" and nitrogen 70 into nonplated gun-bore surfaces 
are discussed in Chapter 14. In the case of carbon, 
the amount does increase in the thermally altered 
layer. Presumably the amount of excess carbon is 
not sufficient to show r an appreciable difference be¬ 
tween the tempered altered layer and the unaltered 
steel. 

Metallographic evidence shows a sharp difference 


CONFIDENTIAL 









246 


THE PRODUCTS OF GUN EROSION 


between the thermally altered layer and the unaltered 
steel, which is usually sorbitic. Since the chemical 
composition is thought to be the same this must be 
only a textural difference. The thermally altered layer 
appears to be structureless when light nital or picral 
etches are used. When, under certain firing condi¬ 
tions, the layer becomes somewhat tempered, it etches 
dark with these reagents, as can be seen in Figure 3. 
With stronger reagents, such as copper potassium 
chloride, it also etches dark and etches even deeper 
than the unaltered steel. 98 Furthermore, it can be 
seen to have a very fine-grained texture. 

The evidence cited above together with the fact 
that the thermally altered layer is harder than the 



Figure 3. A thermally altered layer that has been 
tempered during firing, and therefore etches dark with 
picral. A—thermally altered layer; B—unaltered steel. 
[Cross section of 5-in./51-cal. gun liner No. 806L2 in the 
bag ammunition chamber. Etched with picral (1000X).] 
(This figure has appeared as Figure 20c in NDRC Re¬ 
port No. A-440.) 

sorbitic gun steel have led to the conclusion that the 
layer is martensite although the acicular structure 
characteristic of martensitp is not usually observed. 
It should be mentioned here that needles of some 
phase have sometimes been observed at the junction 
of the thermally altered layer and the inner white 
layer 112 - 261 as shown in Figure 4. They have not as yet 
been identified and can be attributed to iron nitride 
as well as to martensite. Martensite has never been 
identified by x-ray analysis of eroded bore surfaces, 
but this is not surprising. 

Martensite is a term used by x-ray crystallogra- 
phers to designate the tetragonal phase of iron con¬ 
taining dissolved carbon. The axial ratio of the iron 
varies with *the carbon content. When the carbon is 
low the crystal structure is not called tetragonal but 
is usually referred to as distorted cubic. 49 Neverthe¬ 


less, it is still considered martensite as long as doublets 
appear on the x-ray film. 

Metallographers, on the other hand, originally 
used the term martensite in a much less restricted 
sense. For this reason it is probably permissible to 
call this altered layer in guns martensitic, especially 
in view of the fact that even in the case of a slightly 
tempered martensite layer, the texture and the hard¬ 
ness of the layer would be such that it would obvi¬ 
ously be called martensite. Nevertheless, the term 
thermally altered layer is preferred until more is 
known about the nature of the layer. The great im¬ 
portance of this layer in relation to the causes of gun 
erosion is discussed thoroughly in Section 13.2. 



Figure 4. An acicular structure (A) at the junction of 
the thermally altered layer and white layer. A similar 
structure is seen also within the white layer. [Cross sec¬ 
tion of 3.7-in. gun tube No. L/2675 at 21 in. in front of 
the location of the cartridge case. Etched with picral 
{1000X).] (This figure has appeared as Figure 18c in 
NDRC Report No. A-440.) 

It is important to keep in mind that one should not 
expect to find tetragonal martensite in gun-bore sur¬ 
faces unless sufficient carbon had penetrated into the 
thermally altered layer to bring the carbon content 
of this layer above 0.6%. c Iron containing less than 
this amount of carbon might form tetragonal marten¬ 
site on being quenched but this would be self-temper¬ 
ing and would transform to cubic martensite, which 
would appear to be no different from ferrite except 
that the lines of the pattern would be broad as though 
the ferrite were badly strained or not well crystal¬ 
lized. Such broadening of the ferrite lines has some¬ 
times been found. 75 - 123 - 137 


c In the vicinity of resorbing grains of carbide such a con¬ 
centration might be reached. (Informal communication from 
H. E. Merwin, June 26, 1946.) 


CONFIDENTIAL 






SURFACE LAYERS IN ERODED GUNS 


247 


Troostite Band 

The troostite band is a very narrow band which 
forms the transition zone between the thermally al¬ 
tered layer and the unaltered gun steel. It is not 
always sharply defined. Troostite, which is a mixture 
of iron carbide and alpha-iron in extremely fine dis¬ 
persion, may be formed by tempering martensite. 
Since it is a product of thermal transformation, it is 
really a part of the thermally altered layer. It is de¬ 
scribed separately because sometimes it is sharply 
defined due to the fact that it etches darker than the 
rest of the thermally altered layer. 

Inner White Layer 

The inner white layer is in some respects more 
closely related to the thermally altered layer than to 
the outer white layer. When deep etching is applied 
and the specimens are viewed under low magnifica¬ 
tions the inner white layer is usually sharply defined 
because it does not etch. When high magnifications 
are employed, however, the inner white layer and the 
thermally altered layer are usually seen to grade into 
each other. In badly cracked and pebbled surfaces 
the thermally altered layer does not have a uniform 
thickness but extends to a uniform depth below the 
general surface, whereas the thin inner white layer 
actually has a somewhat uniform thickness and fol¬ 
lows the contour of the steel around the edges of the 
cracks. 

Usually the inner white layer shows no structure. 
Sometimes, however, the outermost portion of this 
layer has the structure usually attributed to austen¬ 
ite, 112 - 265 as shown in Figure 5. It would seem that, 
although the thermally altered layer does not take up 
enough carbon or nitrogen to retain austenite on cool¬ 
ing or even tetragonal martensite, yet the outermost 
surface of the original steel can dissolve enough car¬ 
bon or nitrogen to retain austenite. Such an amount of 
carbon or nitrogen would appreciably lower the melt¬ 
ing point of steel. It is believed by some that there is 
sufficient evidence to show that this surface layer had 
melted although it had not been transported. 112 In at 
least one case it was shown that the inner white layer 
was not radically different in chemical composition 
from the steel for it could be tempered to have the 
same type of structure as the steel. 261 Probably, there¬ 
fore, only a portion of the layer was liquefied at the 
peak temperature, the phases present being austenite 
plus melt. Austenite has been consistently identified 


in the bore surfaces of eroded guns by x-ray analysis. 
It may be that this retained austenite is actually the 
dominant constituent of tlje inner white layer. It 
should be emphasized that it is possible to identify 
phases by x-ray analysis that cannot always be read¬ 
ily detected metallographically. 



Figure 5. Grain boundaries in an inner white layer. 

A—inner white layer; B—outer white layer. [Cross sec¬ 
tion of 5-in./25-cal. gun tube No. 1982T at the origin of 
bore. Etched with picral {2000X).] (This figure has ap¬ 
peared as Figure 16c in NDRC Report No. A-440.) 

Outer White Layer 

Outer white layers of the type described here are 
found only in guns that have been fired with single¬ 
base powders. It is true that white layers are some¬ 
times found in guns that have been fired with double¬ 
base powders and two examples of them are shown 
in Figures 6 and 7. Their characteristics are some¬ 
what different from the ones about to be described 
and, unfortunately, they have not been studied in 
sufficient detail to warrant description here. 

The fact that an outer white layer is not found in 
all guns that have been studied does not necessarily 
imply that it was never present, for sometimes traces 
of it can be found in protected areas such as between 
the lips of cracks, as shown in Figure 6. This layer in 
the case of single-base powders is hard, brittle, spalls 
readily, and can thus be easily removed by projectiles 
during firing. For the same reasons, faulty technique 
in polishing the specimens for metallographic exam¬ 
ination removes it. It is not usually found in guns 
where general melting 124 of the surface has occurred. 
With single-base powder, it would probably be found 
somewhat forward of the region of general melting, 
and also on the melted surface. 


CONFIDENTIAL 




248 


THE PRODUCTS OF GUN EROSION 


When deep etches and high magnifications are 
used, a considerable amount of detail can be seen in 
the outer white layer. It is rarely structureless. Blade¬ 
like forms, such as are shown in Figure 8, are ob¬ 
served, which are similar to those seen in cast irons. 112 
In addition, flow structure can be seen. The thickness 
of this layer is variable; it bridges cracks and is thick¬ 
est where it fills up the mouths of cracks, as seen in 
Figure 2. In these places it can be seen to be complex 
as though successive flows had piled up on top of one 
another. 

The boundary between the inner and the outer 
white layers is usually, though not always, sharply 
defined. The band which separates the layers is some¬ 
times seen to consist of some extraneous substance, 
perhaps a decomposition product of one of the white 



Figure 6. An outer white layer in the mouth of a crack 
in a gun that had been fired with double-base powder. 
[Longitudinal cross section of 37-mm gun tube No. 
T-13 in the region Yi. to 1^2 in. from the origin of bore. 
Etched with picral (1500X).] (This figure has appeared 
as Figure 18d in NDRC Report No. A-440.) 

layers. 261 Some extraneous material, including copper, 
is often found throughout the outer white layers. 112 
This feature is illustrated in Figure 9. d 

There is no doubt in view of all this evidence that 
the outer white layers have been at least partially 
liquefied and flowed into place. (For other evidences 
of liquefaction of the bore surface, see Section 10.5.2.) 
There is also no doubt that, whereas the lower altered 
layers can still be considered steel in various textural 
modifications, the outer white layers are of quite a 
different chemical composition than the underlying 
steel, for experiments have shown that this outer 
zone does not change on tempering. 261 It is concluded 


d The copper is even more obvious in a Kodachrome 112 of 
the same field illustrated in the figure. 


that compounds of iron that are formed by the reac¬ 
tion of powder gases on gun steel are in some manner 
related to the outer white layers. In recent studies at 
the Geophysical Laboratory of the products of gun 
erosion 98 considerable emphasis was placed on finding- 
methods to determine the chemical composition of 
outer white layers and to identify the constituent 
phases. 



Figure 7. An outer white layer in a gun that had been 
fired with double-base powder. [Cross section of 37-mm 
gun tube No. 43928 at 1 7 /g in. in front of the location of 
the cartridge case. Etched with picral (1500X).] (This 
figure has appeared as Figure 18e in NDRC Report 
No. A-440.) 


12 2 DIFFERENCES BETWEEN GUNS FIRED 
WITH SINGLE-BASE AND WITH 
DOUBLE-BASE POWDERS 

12 21 X-Ray Examinations of 

Eroded Bore Surfaces 

Compounds of iron on the bore surface were first 
identified by means of x-ray analysis of eroded bore 
surfaces of Service guns. In the earliest work only fer¬ 
rite and austenite 6 were found. 483 Later work took 
cognizance of unidentified lines in the x-ray patterns, 
in addition to the two modifications of iron contain- 


e Ferrite is body centered cubic iron containing carbon, and 
austenite is face centered cubic iron containing carbon, but in 
both cases nitrogen can replace some of the carbon. Pure iron 
undergoes a phase change at 910 C. Below this temperature it 
is body centered cubic and above this temperature face 
centered cubic. These two phases of pure iron are called alpha 
and gamma respectively. These terms are sometimes used by 
the x-ray crystallographers to designate the structure of the 
iron even when not pure; however, in this case, the names 
ferrite and austenite are preferred. 


CONFIDENTIAL 







EROSION BY SINGLE-BASE AND DOUBLE-BASE POWDERS 


249 


ing carbon or nitrogen or both. It was thought that 
iron nitrides could not account for these lines but the 
possibility that they might be due to a complex iron 
cyanide was suggested. 261 

Compounds of iron have recently been identified 
in eroded bore surfaces for the first time. 28 - 49 All the 
examinations were made at or near the origin of ri¬ 
fling. In addition to ferrite and austenite the follow¬ 
ing compounds were found: wiistite (FeO), cementite 
(Fe 3 C), and the epsilon phase of iron nitride (Fe 2 N a: ). 
The first was found in relatively large amounts in 
guns fired with double-base powder and sometimes in 
minor amounts in those fired with single-base powder. 



Figure 8. The blades of cementite in an outer white 
layer composed of several superposed flows. [Cross sec¬ 
tion of 3-in. gun liner No. 1460 at in. in front of the 
location of the cartridge case. Etched with picral 
{750X).\ (This figure has appeared as Figure 12a in 
NDRC Report No. A-440.) 

The last two were found only in guns that had been 
fired with single-base powder. 

Examples of guns in which wiistite was the only 
compound identified were a 3.7-in. Canadian gun 28 
and a 37-mm M3 gun tube. 50 Examples of guns rich 
in cementite and epsilon iron nitride were a 90-mm 
gun tube 49 and the 3-in. gun liner No. 1460. 98 A 5-in./ 
25-cal. gun tube, 28 - 50 which had been fired with Pyro 
powder, was found to contain the products character¬ 
istic of both single-base and double-base powders. 

12 2 2 X-Ray and Electron Diffraction 
Examinations of Test Blocks 

The above differences between the guns fired with 
single-base and with double-base powders are in 


agreement with the results obtained on test blocks 
that were subjected to explosions of powders of these 
types in the apparatus described in Section 11.2.4. 31 

It was found that a specimen of gun steel that had 
been exposed to the gases from double-base powder 
had a layer of wiistite about 0.01 mm thick on the 
surface. The film on the surface of a specimen that 
had been exposed to the gases from single-base pow¬ 
der, on the other hand, contained no wiistite but in¬ 
stead consisted largely of cementite. This layer of 



Figure 9. Patches of copper within an outer white 
layer. A—inner white layer; B—outer white layer; 

C—patches of copper. [Cross section of 5-in./25-cal. 
gun tube No. 1982T at the origin of bore. Etched with 
picral ( 750X ).] (This figure has appeared as Figure 
16a in NDRC Report No. A-440.) 

cementite, which in one case was about 0.001 mm 
thick, was overlaid by a layer of a complex iron cy¬ 
anide about one-tenth as thick. The cementite layer 
was removable in large flakes by means of nital, and 
beneath it was found a considerable amount of addi¬ 
tional cementite and still larger amounts of austenite. 

12,2,3 X-Ray Examinations of 

Fired Metal Particles 

The results of the examinations of guns were also 
in accord with those obtained by examining steel fil¬ 
ings that had been mixed with the powder charge in 
a caliber .30 rifle, 28 and fired according to the tech¬ 
nique described in Section 11.2.6. Wiistite formed 
about 90% of the products that resulted from the fir- 


CONFIDENTIAL 




250 


THE PRODUCTS OF GUN EROSION 


ing of electrolytic iron with double-base powder. The 
rest of the material was ferrite together with a small 
amount of austenite. 

Both electrolytic iron and gun steel were fired with 
IMR, the hottest of the single-base powders. The re¬ 
sults were practically the same in both cases. The 
main products were ferrite and austenite. Fair 
amounts of epsilon iron nitride and wiistite were also 
present. No cementite was found. Undoubtedly the 
particles attained too high a temperature for cement¬ 
ite to be the end product. The metal particles were at 
a higher temperature for a longer time than the bore 
surface, which was rapidly cooled by the mass of gun 
steel. 

Iron particles of several different meshes were fired 
with single-base powder. The larger amounts of wus- 
tite and epsilon iron nitride found in the reaction 
product when iron of smaller particle size was fired, 
can be readily accounted for by the increased surface 
area. These results indicate how much more reactive 
a bore surface may be once it has acquired an exten¬ 
sive crack system. 

12 2 4 Thermodynamic Considerations 

The problem of determining the theoretical com¬ 
pounds resulting from the interaction of iron with 
powder gases was studied in order to explain the dif¬ 
ferences observed between the results obtained with 
single-base powders and those obtained with double¬ 
base powders. One means of attack 60 employed stan¬ 
dard thermodynamic methods to determine, under a 
number of conditions, the ultimate product, that is, 
the product that would be formed if chemical equi¬ 
librium were established between a finite quantity of 
iron and an infinite quantity of the powder gas. In a 
competing series of reactions of gases with iron, the 
ultimate product is the one that requires the greatest 
free-energy change for formation directly from iron. 
All possible equilibriums between the gases and iron 
are discussed in Section 13.3.3. It seems improbable 
under gun-firing conditions that the reactions in the 
gas phase are fast enough to establish the equilibrium 
composition of the gas at the temperature of the bore 
surface immediately. Thus the reactions on the sur¬ 
face are those with a gas at the temperature of the 
surface but having the composition it would have at 
a somewhat higher temperature. 

It was found that for FNH-M1 powder (single¬ 
base) the ultimate product at low surface tempera¬ 
tures is Fe 3 0 4 if the reacting gas is cooled slowly to the 


surface temperature, whereas it is Fe 3 C if the gas at 
the surface is cooled rapidly. At higher surface tem¬ 
peratures FeO is formed instead of Fe 3 0 4 . This ac¬ 
counts for the absence of cementite in the products 
formed by firing steel or iron filings, mentioned in the 
preceding subsection. 

For FNH-M2 powder (double-base) the ultimate 
products at low surface temperatures are the same as 
for single-base powder. At high surface temperatures, 
however, the product is FeO for a wide range of con¬ 
ditions of cooling of the reacting gas. If 0.1% sulfur 
is added to either powder, the ultimate product at 
low surface temperatures is FeS. The presence of sul¬ 
fur does not affect the products obtained at high sur¬ 
face temperatures. 

The formation of iron nitrides was not considered 
quantitatively because of the paucity of the data. 
Extrapolation of available data led to the conclusion 
that formation of iron nitrides is impossible at the 
temperatures of bore surfaces during firing. Nitrides, 
however, have been found in guns. It may be that 
they undergo a transition at temperatures higher 
than those at which they have been studied, thus in¬ 
validating any extrapolation of the data. 

In this research the formation of solid solutions 
could not be considered because of the lack of data. 
This was unfortunate because the penetration of car¬ 
bon and nitrogen into the steel, evidence for which is 
given in Chapter 14, does take place presumably with 
the formation of solid solutions. It should be borne in 
mind, however, that once the bore surface has taken 
up enough carbon or nitrogen or both to liquefy, the 
formation of solid solutions is of no concern. The li¬ 
quefied material can dissolve carbon and nitrogen 
more readily than a solid phase such as austenite. 

The other method of attack 138 was an attempt to 
explain the differences between the results found in 
the experiments with the test blocks and those in the 
experiments with the metal filings. Thermodynamic 
arguments were used to determine the stabilities of 
the solid phases and also to predict the direction of 
the chemical reactions. Equilibrium curves were plot¬ 
ted for the reactions to form FeO and Fe 3 C from C0 2 
and CO and for th6 reaction to form carbon and car¬ 
bon dioxide from the monoxide. The curves, shown in 
Figure 8 of Chapter 13, bounded the regions for the 
positive and negative free energies of the reactants 
and the products. The various temperatures consid¬ 
ered were those of the bore surface during firing. In 
this work great care was taken to show that under 
the conditions of gas flow in the experiments with the 


CONFIDENTIAL 



EROSION PRODUCTS ENTRAPPED IN THE COPPERING 


251 


test blocks the assumption that the equilibrium con¬ 
stants to be used for gaseous reactants correspond to 
a temperature only slightly above the bore surface 
temperatures is justified. In the use of these equi¬ 
librium constants rather than those for higher tem¬ 
peratures this second investigation differed from the 
first. 

A consideration of the effects of pressure and tem¬ 
perature upon the reactions involved led to some im¬ 
portant conclusions, which will be discussed in Sec¬ 
tion 13.3.3. Also the effects of nitrogen and hydrogen 
upon the system Fe-C-0 were considered. Low tem¬ 
perature reactions were discussed in this connection. 
Some of the conclusions were borne out by experi¬ 
mental evidence. 

The following conclusions were drawn from the 
study of the Fe-C-0 system. If the C0/C0 2 ratio is 
of the order of 2 or 3 and if the temperature at the 
surface of the specimen does not exceed about 1100 C, 
Fe 3 C is to be expected as the major reaction product, 
but FeO is to be expected above about 1200 C. If the 
CO/C0 2 ratio is of the order of unity, then FeO is to 
be expected as the major reaction product above 
about 1000 C, but below about 900 C Fe 3 C may still 
be produced in considerable quantity. Increasing the 
pressure favors Fe 3 C while decreasing the pressure 
favors oxide formation. The higher CO/C0 2 ratio 
corresponds roughly to single-base powders and the 
lower ratio to double-base powders. 

It should be noted that in this second investigation 
the products were termed major rather then ultimate. 
Nothing is known about the rates of reaction at the 
temperatures and pressures that are obtained in the 
firing of guns, but presumably kinetics are of some 
importance. Therefore, even though all the theoreti¬ 
cal ultimate products have been found in guns, it 
cannot safely be assumed that an actual end product 
will always be an ultimate product. 

12 2 5 X-Ray Examinations of Barrels Fired 
in the Erosion-Testing Gun 

The quantitative determination of the rate of ero¬ 
sion of gun steel with different propellants fired in the 
caliber .50 erosion-testing gun is thoroughly discussed 
in Section 15.3. In addition to the general relations 
between the flame temperature of the powder and the 
erosion rate and thermal transformation of the steel, 
the differences in the compounds that were identified 
on the bore surfaces by x-ray examination are related 
to the erosiveness of the powder. 


The erosiveness depends not only on the flame 
temperature but to some extent on the composition 
of the powder. Ferrite and'austenite were found in 
all cases. With double-base powders, the only other 
constituent identified was wtistite (FeO). Smaller 
amounts of this substance were found for all the 
cooler powders with the exception of IMR, the hot¬ 
test of the single-base powders. Cementite (Fe 3 C) 
and the epsilon phase of iron nitride (Fe 2 N*) in 
varying amounts were revealed by the x-ray analysis 
of the barrels fired with most of the cooler powders. 
In the case of the four coolest powders, an unidenti¬ 
fied constituent was found. 

In view of the thermodynamic considerations men¬ 
tioned above, the wtistite and cementite are to be ex¬ 
pected in those cases where they were found. The 
greater complexity of the results is probably due to 
the fact that other molecules besides those of the wa¬ 
ter gas reaction are present in the propellent gases. 
The presence of nitrogen in some form or other ac¬ 
counts for the nitrides that are found. 

The dependence of erosion rate on the temperature 
of the propellent gases in contact with the bore sur¬ 
face and on the compositions of the powders, a major 
factor in both cases being the CO/C0 2 ratio, has led 
to the theory that a powder of a certain composition 
might be found that would be less erosive than any of 
the present powders. This theory is discussed in Sec¬ 
tion 15.6.3, and several preliminary experiments 
along this line are described in Section 15.6.5. 

123 EROSION PRODUCTS ENTRAPPED 
IN THE COPPERING 98 

12,8,1 Coppering 

The deposits of copper or gilding metal found in 
guns are rarely clean, since they entrap material 
eroded from the bore surface, including compounds 
formed in reactions with the powder gases. For this 
reason heavily coppered sections of eroded guns have 
been used as a source of erosion products for study. 
These compounds were segregated from the copper 
by dissolving the latter by the technique described in 
Section 11.4.3. As was stated in Section 10.5.4, the 
heaviest coppering commonly occurs in the central 
portion of the bore, thus the deposits that were stud¬ 
ied yielded a wealth of material that had been carried 
forward from the region of the origin of rifling. Fre¬ 
quently a pronounced layered structure was evident, 
consisting of alternate deposits of reaction products 


CONFIDENTIAL 




252 


THE PRODUCTS OF GUN EROSION 


and other debris and of flows of copper. Because of 
this mode of deposition, which was likened to alter¬ 
nate ash falls and lava flows from a volcano, it was 
possible to discern whether or not there was any dif¬ 
ference between early and later deposits. 

12 3 2 Sulfur in the Coppering 

In the study of gun erosion it is the reactions of the 
powder gases with steel that are of interest; however, 
the study of the coppering in guns has yielded infor¬ 
mation regarding other reactions that have taken 
place. One of these is the reaction of the copper or 
gilding metal with the sulfur in the powder gases. The 
fact that copper and zinc react so readily with the 
sulfur may mitigate the effect of the sulfur compon¬ 
ents of the powder on the gun steel, which is discussed 
in Section 14.4. 

Copper Sulfides 

Both the sulfur which is dissolved in the copper 
and that which is present in the form of copper sul¬ 
fides can easily be detected without removing the 
copper deposit. When coppered bore-surface speci¬ 
mens are heated in hydrogen and the effluent gases 
are passed into cadmium acetate, a yellow precipitate 
of cadmium sulfide is formed which indicates the 
presence of hydrogen sulfide. The presence of copper 
sulfides can also be indicated by a contact print 
method in which potassium cyanide is used to etch 
the specimen. Copper sulfides, such as chalcocite 
(Cu 2 S) and covellite (CuS), are soluble in potassium 
cyanide. The total sulfur in the copper can be found 
by removing the deposit with the ammoniacal solu¬ 
tion described in Section 11.4.3, and by analyzing 
both the solutions and the insoluble residues. The 
solution dissolves the chalcocite and covellite as well 
as copper; the relatively insoluble copper sulfide, 
digenite (CugS 5 ), can sometimes be found in the in¬ 
soluble residue by x-ray analysis. 

Chalcocite has been identified by this method of 
analysis in the material removed from the cracks 
near the origin of rifling where there was no copper 
obvious on the bore surface. 

Zinc Sulfides 

When gilding-metal (alloy of copper with 10% 
zinc) rotating bands have been used, zinc sulfides are 
found in the coppering. These cannot be detected by 


contact printing methods but can readily be identi¬ 
fied by x-ray analysis of residues centrifuged from 
ammoniacal decoppering solutions (Section 11.4.3), 
for they are less soluble than the copper sulfides. Both 
wurtzite (hexagonal ZnS) and sphalerite (cubic ZnS) 
have been found in eroded guns. In the case of a 
75-mm gun a very interesting succession of deposits 
was found interlayered with the coppering. Wurtzite 
alone was found in the deposit at the surface, while 
both sphalerite and wurtzite and also digenite were 
found nearer the copper-steel interface. Wurtzite and 
sphalerite are the high- and low-temperature forms of 
zinc sulfide respectively, the inversion point with the 
pure sulfide is 1020 C. Digenite (Cu 9 S 5 ) is a low-tem¬ 
perature form of copper sulfide. Presumably, this in¬ 
version temperature would enable one to tell what 
temperatures were reached at and near the bore sur¬ 
face, but the relationships among these sulfides are 
not simple. The absence of sphalerite does not nec¬ 
essarily mean that the immediate bore surface had 
reached a temperature of 1020 C since the inversion 
point may be lowered by the presence of other ele¬ 
ments, especially iron, dissolved in the zinc sulfide. 484 

Zinc sulfides are sometimes found in guns that have 
copper rather than gilding-metal deposits. The zinc 
in these cases may have been derived from impurities 
in the copper or what is more likely from the car¬ 
tridge case. An example of the latter is undoubtedly 
the wurtzite that was found at the location of the 
mouth of the cartridge case in a 5-in. gun. 28 

Other Sulfur Compounds 

Galena (PbS) has been found in the copper and 
also on bore surfaces where there was no copper. The 
lead may be derived from the primers or from lead 
foil incorporated in the powder charges for the pur¬ 
pose of decreasing the amount of coppering. Metallic 
lead has been found in eroded guns. 123 - 137 Barium sul¬ 
fate and carbonate have been identified in residues 
centrifuged from decoppering solutions. Potassium 
sulfate has likewise been identified in bore-surface 
products. The barium and potassium were both de¬ 
rived from the powders. 

Distribution of Sulfur in the Coppering 

Very little study has been made of the distribution 
of sulfur in the coppering of guns. In only one gun, a 
5-in. Naval gun tube, have the relations between 
the amounts of sulfur and copper with respect to the 


CONFIDENTIAL 



EROSION PRODUCTS ENTRAPPED IN THE COPPERING 


253 


distance from the origin of rifling been studied. The 
results are given in Table 1. It can be seen that the 


Table 1 . Copper and sulfur on the bore surface of a 5-in. 
Naval gun at different distances from the origin of 
rifling. 98 


Distance from 
origin of rifling 

Copper Sulfur 


(in.) 

(mgs/sq cm) (mgs/sq cm) 

S/Cu 


1.4- 3.0 

6 

0.11 

0.019 

3.6- 5.2 

15 

0.24 

0.016 

5.2- 6.8 

30 

0.15 

0.005 

11.2-12.8 

42 

0.25 

0.006 

11.5-12.8 

41 

0.20 

0.005 

13.8-18.8 

41 

0.12 

0.003 


amount of copper increased with distance from the 
origin of rifling, reached a maximum at about 13 in., 
and apparently decreased slowly from there on. The 
amounts of sulfur were erratic, but the ratios of sulfur 
to copper seem to indicate that the sulfur in the pow¬ 
der gases was gradually being exhausted by combina¬ 
tion with the copper. 

12 3 3 True Erosion Products 

Entrapped in the Copper 

True erosion products are compounds of iron or of 
any of the alloying constituents in the gun steel. 
These, as well as the sulfur compounds mentioned 
above, have been found entrapped in the coppering 
of guns. 

Cementite (Fe 3 C) was identified in the decoppering 
residues from all but one of the guns which had been 
fired with single-base powders. In the case of the one 
exception, a 14-in. Naval gun, it may well be that the 
region studied was too near the muzzle end. In Sec¬ 
tion 12.2.1 it was pointed out that cementite is a 
characteristic product at and near the origin of rifling 
in guns that have been fired with single-base powder. 

X-ray analyses of the bore surface had also shown 
that the epsilon phase of iron nitride (Fe 2 N ar ) was 
characteristic of all these guns and that small amounts 
of wlistite (FeO) were found on some. This iron ni¬ 
tride is unusually resistant, even more so than ce¬ 
mentite, to attack by the ammoniacal decoppering 
solution; thus if present in the copper, it should be 
found in the residues obtained on decoppering. Epsi¬ 
lon iron nitride was identified in some of the residues 
in which cementite had been found but was not de¬ 
tected in those from the guns in which wlistite had 
also been found on the bore surface. This may mean 


that the iron nitride is less stable than cementite 
when they are formed under conditions that are fa¬ 
vorable for the formation of a small amount of wiis- 
tite; and that, while iron nitride is found on bore 
surfaces together with both wlistite and cementite, 
it may not be stable when carried away from the bore 
surface and entrapped in the copper. As was implied 
in Section 12.2.1 on thermodynamic considerations, 
the question of the formation and stability of the iron 
nitrides in guns demands further investigation. 

Austenite, as is stated in Section 12.2.1 and 15.3.4, 
has been found in the bore-surface layers of all eroded 
guns that were examined by x-ray diffraction, no 
matter what type of powder had been used. It has 
not been found in residues obtained on decoppering. 
As a matter of fact, if it were found, it would be nec¬ 
essary to evaluate the results with great caution since 
even small amounts of copper contaminating the resi¬ 
due would give the same x-ray diffraction pattern as 
austenite. 

Wlistite (FeO), which is the only iron compound 
that was found by x-ray diffraction studies of guns 
fired with double-base powder (see Sections 12.2.1 
and 15.3.4), is also the only iron compound identified 
in decoppering residues obtained from such guns. 

Other iron compounds that were not found directly 
on the bore surfaces have been found entrapped in 
the copper. One of the most interesting of these w^s 
iron-rich, brown “enamel,” which was the most out¬ 
standing feature of a 5-in. chromium-plated Naval 
gun. This contained a small amount of crystalline 
material that could be identified by x-ray diffraction 
analysis. The compounds identified were potassium 
copper ferrocyanide K 2 CuFe(CN)6, wlistite, barium 
sulfate, and sphalerite (cubic ZnS). The complex cy¬ 
anide is, in a sense, a true erosion product for it con¬ 
tains iron. It probably represents the reaction be¬ 
tween the copper and a complex iron cyanide formed 
on the bore surface of the gun. Other indications of 
complex iron cyanides on surfaces of guns and test 
specimens have been found. 31 ’ 98 ’ 261 

Wlistite was identified directly on the bore surface 
of this gun but the area subjected to the x-ray beam 
is known to have contained a considerable amount of 
brown enamel and the wlistite is believed to be closely 
associated with the brown enamel. Wurtzite (hex¬ 
agonal ZnS) was found at the location of the mouth 
of the cartridge case, thus it is not surprising that 
sphalerite (cubic ZnS) was detected beyond the origin 
of rifling. These two positions in the bore of this gun 
are shown in Figures 10 and 11, respectively. 


CONFIDENTIAL 







254 


THE PRODUCTS OF GUN EROSION 


Chemical analysis of the brown enamel revealed its 
complexity. A fair amount of carbonaceous material 
was present. Although the enamel was not homogene¬ 
ous, the following analysis can be given as representa¬ 
tive of the ignited material. The constituents are listed 
on the basis that the material is a silicate: Si0 2 5.8%; 
BaO 1.9%; PbO 7.8%; CuO 11.2%; Fe 2 0 3 64.4% ; 
Cr 2 0 3 absent; ZnO 5.5%; Mn 3 0 4 0.8%. The sulfur 
content, which must be referred to the barium sulfate 
and zinc sulfide, identified by the x-ray analysis, was 
also determined. Part of the iron must be accounted 
for by the wiistite and complex iron cyanide, but the 



Figure 10. Surface of the 5-in./25-cal. gun tube No. 
1982 T at the location of the mouth of the cartridge 
case. ( 12X .) (This figure has appeared as Figure 8b in 
NDRC Report No. A-440.) 

bulk of it was probably in the ferrous condition in the 
amorphous glassy material. It is believed that the 
wiistite associated with the brown enamel was not 
derived directly from the interaction of the powder 
gases with the gun steel but was crystallized out of the 
iron-rich, brown enamel. How the iron became incor¬ 
porated in this enamel is not known. f 

Pyrrhotite (FeS + S) was identified in the de- 
coppering residue from only one gun. The fact that it 
is seldom found agrees with the results of the thermo¬ 
dynamic calculations which showed that ferrous sul¬ 
fide is not often likely to be found in eroded guns, 
since its formation involves a self-inhibiting reaction. 60 

f For further details concerning this interesting product the 
reader should consult the original work. 98 


12,3,4 Complexity of Decoppered Residues 

It has been shown above that the brown enamel is 
a complex substance. This complexity also applies to 
most of the residues obtained on decoppering. It 
should be remembered that x-ray analysis reveals the 
presence of crystalline species but does not reveal 
their chemical complexity. Furthermore, the applica¬ 
tion of x-ray analysis® was further limited by the fact 
that many unidentified lines were present in the pow¬ 
der photographs, and it was obvious in some cases 
that considerable amounts of amorphous materials 
were present which made it difficult to get good, clear 
patterns. In addition, chemical analyses, other than 
the ones of the brown enamel, were made which 
showed large amounts of carbonaceous matter in 
some cases and also minor amounts of other sub¬ 
stances some of which may have been in solid solution 
in the compounds which were identified. 

12 4 EROSION PRODUCTS SEGREGATED 
BY MECHANICAL MEANS 98 

12,4,1 Erosion Products Collected in Cracks 

The material filling the deep cracks of an eroded 
3-in. gun was removed mechanically by the method 
described in Section 11.4.2 and was subjected to x-ray 
and chemical analysis. Austenite and magnetite 
(Fe 3 0 4 ) were identified in addition to sulfides of cop¬ 
per and zinc. Chemical analysis showed that 4.7% of 
carbon and 0.11% of nitrogen were present. 

Since cementite was not detected by x-ray diffrac¬ 
tion, the presence of extraneous carbonaceous mate¬ 
rial must be considered to account for this large 
amount of carbon. In view of the fact that guns are 
greased, it is more than likely that this grease, lodged 
in the major crack system, had been repeatedly sub¬ 
jected to hot gases, had thereby undergone destruc¬ 
tive distillation, and had left a coherent coke-like 
substance in the cracks. It is obvious that if chemical 
analysis only is used in determining the presence of 
carbon, no inferences can be drawn with respect to 
the presence or absence of cementite. The nitrogen 
content was not large enough to permit the deter¬ 
mination of either of the nitrides of iron by x-ray 
analysis. Moreover, this small amount of nitrogen 
may have been dissolved in the austenite and thus 
escaped detection by this method of analysis. 

g The present limitations in x-ray technique are discussed at 
the end of Section 11.5.2. 


CONFIDENTIAL 






CONSTITUENTS OF SURFACE LAYERS REMOVED CHEMICALLY 


255 


The thermochemical calculation referred to in Sec¬ 
tion 12.2.4 indicated that magnetite (Fe 3 0 4 ) would be 
the ultimate product resulting from the interaction 
of the powder gases and the steel at lower tempera¬ 
tures than are attained at the bore surface during 
firing. Thus, while cementite is formed on the bore 



% 

v..- • 

db-> ; 


Figure 11 . A crack filled with a “slag’’ material that 
may be the same as the “brown enamel” disengaged 
from the bore surface of the same gun. A—outer white 
layer; B —slag-filled crack. [Cross section of 5-in./25- 
cal. gun tube No. 1982T at origin of bore. Etched with 
picral ( 500X ). (This figure has appeared as Figure 15d 
in NDRC Report No. A-440.) 


surface, magnetite will be formed in the deep cracks 
which extend well below this surface. 

12,4,2 Constituents of Surface Layers 
Removed Mechanically 

The outer white layer, which was described in 
Section 12.1.2, spalls readily and in some cases it is 
possible to remove portions of this layer mechanically. 
X-ray analysis of some of this material from a 3-in. 
gun liner revealed the presence of cementite, austen¬ 
ite, and epsilon iron nitride, the same three products 
that had been found by examinations of the bore sur¬ 
face, which indicates that the compounds identified 
on the eroded bore surfaces of guns are definitely con¬ 
stituents of at least the outer white layer. In the case 
of a 57-mm gun, the layer removed was found to con¬ 
tain cementite and the gamma-prime phase of iron 
nitride (Fe 4 N), whereas the epsilon iron nitride 
(Fe 2 N x ) together with cementite had been found in 
the decoppering residues. 


12 5 CONSTITUENTS OF SURFACE LAYERS 
REMOVED CHEMICALLY 

12,5,1 Introduction 

The resistance of the white layers to the usual, 
mild, etching reagents was discussed in Section 12.1.2. 
Etching with boiling alkaline sodium picrate, how¬ 
ever, showed a concentration of cementite in the 
outer white layer. 112 In Figure 12 is shown the ap¬ 
pearance of this outer layer after etching. The high 
concentration of cementite is indicated by the intense 
blackening of the surface portion. This blackening 



Figure 12. The relative concentrations of cementite 
in an outer white layer as shown by etching with boiling 
alkaline sodium picrate. A—thin surface film that has 
etched entirely black; B —steel. [Cross section of 3-in. 
gun liner No. 881 at 12 in. in front of location of 
cartridge case. Etched with boiling alkaline sodium 
picrate (S50X).] (This figure has appeared as Figure 
12e in NDRC Report No. A-440.) 


decreases in intensity below the surface and thus in¬ 
dicates a decrease in concentration of cementite. 

Since a solution of copper potassium chloride is 
useful in segregating cementite from gun steel, as de¬ 
scribed in Section 11.4.4, the use of this reagent to 
segregate the white layers from the underlying steel 
was undertaken. That this might be done was first 
demonstrated by etching a specimen typical of an 
eroded gun that had a cementite-rich surface layer. 98 
Both the thermally altered layer and the unaltered 
steel were deeply etched, the former more so than the 
latter, while the white layers were but little attacked. 


CONFIDENTIAL 




256 


THE PRODUCTS OF GUN EROSION 


Interstitial material had been dissolved out of the 
outer white layer leaving the blades of cementite 
sharply defined. When eroded bore surfaces had been 
subjected to a solution of copper potassium chloride, 
coherent flakes could be removed. Metallographic ex¬ 
amination showed that the bulk of the material re¬ 
moved in this manner was outer white layer from 
which the interstitial material had been dissolved by 
the solution. 112 One of these flakes is shown in Figure 
13. By a refinement of this technique, it was possible 
to obtain casts of the crack systems of eroded guns 
by the method described in Section 11.4.4. 



Figure 13. Cross section of outer white layer removed 
by chemical means. [3-in. gun liner No. 1460 at 3^ in. 
in front of the location of the cartridge case. Not etched 
(500X).] (This figure has appeared as Figure 17 in 
NDRC Report No. A-440.) 

12,5,2 Study of Cementite-Rich 

Surface Layers 

An extensive study was carried out on flakes re¬ 
moved from a 3-in. gun liner that was typical of guns 
fired with single-base powder. 98 A moderately thick 
deposit of flows of outer white layer material had 
spread out on the bore surface in the area at and near 
the origin of rifling. 112 These complex flows are shown 
in Figure 8. X-ray analysis of the flakes revealed that 
the dominant constituent was cementite (Fe 3 C). In 
addition chromium carbide (Cr 7 C 3 ) and the gamma- 
prime phase of iron nitride (Fe 4 N) were found in the 
material from the lands and a lesser amount of the 
latter in that from the grooves. Chemical analyses of 
the flakes bore out the x-ray results. The quantitative 
analyses of the material chemically removed from the 
grooves showed that it was essentially iron carbide as 
truly represented by the formula (Fe 3 C) usually 
given for cementite. The purity of this cementite con¬ 
trasted sharply with that of the cementite segregated 
from the steel. The analyses are given in Table 2. In 


Table 2. Analysis of 3-in. gun steel and of carbides 
segregated from the steel and from the bore surface. 98 



Gun steel 
(%) 

Carbide from 
steel 
(%) 

Carbide from 
surface* 

(%) 

Fe 

97.35 f 

64.67 

87.0 

Cr 

0.90 

13.05 

0.7 

Mn 

0.74 

4.81 

0.2 

Mo 

0.47 

5.03 

0.2 

V 

0.07 

2.2 


Cu 


0.56 

None 

P 

0.009 

trace 


S 

0.019 

0.59 


Si 

0.26 

0.091 

None 

C 

0.17 

6.80 

6.8 

N 

0.007 

0.34 

0.3 


* Average of several samples each of which contained flakes from several 
grooves. t By difference. % Determined as Si02. 


fact, the cementite from the groove surfaces was ac¬ 
tually richer in iron with respect to alloying constitu¬ 
ents than was the steel itself, as can be seen from the 
ratios given in Table 3. 


Table 3. Ratios of alloying constituents to iron for gun 
steel and for cementite removed from the groove sur¬ 
faces of a 3-in. gun. 98 



Gun steel 

Cementite 

Cr/Fe 

0.0093 

0.0081 

Mn/Fe 

0.0076 

0.0023 

Mo/Fe 

0.0048 

0.0023 


Cementite, of course, had been identified on the 
bore surface of this gun (Section 12.2.1) and also in 
the outer white layer that had been removed mechan¬ 
ically (Section 12.4.2). Austenite and epsilon iron 
nitride (Fe 2 N x ) had also been detected in the outer 
white layer (Section 12.4.2). The austenite was prob¬ 
ably a constituent of the interstitial material that 
had been dissolved during the segregation. It is not 
surprising that the epsilon iron nitride was not found 
in the material that was removed from the surface 
with copper potassium chloride. Experiments showed 
that this nitride is less resistant to attack by this 
solution than the gamma-prime phase. This very fact 
permitted the concentration of the latter so that it 
could be identified by x-ray diffraction. The exact 
location of the iron nitrides in the outer white layer is 
not known. 

The chromium carbide that was identified in the 
material from the land surfaces was not detected by 
direct examination of the bore surface, and, like the 
gamma-prime iron nitride, was found by virtue of its 


CONFIDENTIAL 
















CONSTITUENTS OF SURFACE LAYERS REMOVED CHEMICALLY 


257 


being concentrated due to the removal of other sub¬ 
stances by the copper potassium chloride solution. 
This result is not unexpected, since the gun steel con¬ 
tains chromium which is a good carbide-former. Un¬ 
der certain conditions chromium carbide is formed 
more readily than iron carbide. This chromium car¬ 
bide was also found at the muzzle end of a 5-in., 



Figure 14. A thick white layer which has a well-de¬ 
fined structure only in the outermost portion. A —outer¬ 
most outer white layer; B— innermost outer white layer; 

C—inner white layer, which has etched dark with this 
rather strong etch; D—thermally altered layer. [Cross 
section of 76-mm gun tube M1A1 No. 1425 at the loca¬ 
tion of the mouth of the cartridge case. Etched with 
picral HC1 ( 1000X ).] (This figure has appeared as 
Figure 13d in NDRC Report No. A-440.) 

chromium-plated gun where the plate was essentially 
still intact except along the driving edges of the 
lands. 98 The presence of this carbide was revealed 
by x-ray analysis of the crack material segregated 
by the technique of Vinylite plaques described in 
Section 11.4.4. 

Cementite has been found to be the dominant con¬ 
stituent of surface layer flakes from other guns that 
had been fired with single-base powders. 98 

12 5 3 Study of Austenite-Rich Surface Layer 

A 76-mm gun tube that was studied 98 ’ 112 ’ 265 was 
found to have a different type of outer layer than the 
one described above. The steel of this tube contained 


4.30% nickel. In front of the location of the mouth of 
the cartridge case the white layer was unusually thick 
but forward of this area it was present as a very thin 
film. Where the layer was thinly spread over the sur¬ 
face, the bore displayed all the features associated 
with general melting of the bore surface as shown in 
Figure 17 of Chapter 10. According to metallographic 
and x-ray evidence, the thick portion of the white 
layer was not very different from the white layers in 
other guns that were studied. The outermost por¬ 
tion of the outer white layer had the bladelike forms, 
which are believed to be the cementite that was iden- 



_ w. m . 

c 


Figure 15. A seemingly structureless, thin white layer. 

A—white layer; B—steel; C—nonmetallic inclusion; 

D—pocket formed by the fluxing of an inclusion at the 
bore surface. [Cross section of 76-mm gun tube M1A1 
No. 1425 at 14^6 in. from the origin of rifling. Etched 
with picral (2000X).] (This figure has appeared as 
Figure 2b in NDRC Report No. A-440.) 

tified by x-ray diffraction. The inner portion of the 
outer white layer seemed structureless. These fea¬ 
tures are shown in Figure 14. Nickel carbide (Ni 3 C) 
was identified on the surface in addition to cementite. 
The thin white layer, Figure 15, in the smooth, rip¬ 
pled area, etched a brownish gray but no well-defined 
structure was apparent. In this region no cementite 
was detected on the bore surface, but nickel carbide 
was found here together with austenite. 

This thin white layer was removed from the bore 
surface by means of copper potassium chloride solu¬ 
tions. X-ray and chemical analysis showed that the 
surface layer flakes contained nickel carbide, austen¬ 
ite, approximately 30% of carbon, and a much smal¬ 
ler percentage of nitrogen. There was possibly some 


CONFIDENTIAL 






258 


THE PRODUCTS OF GUN EROSION 


gamma-prime iron nitride present, but probably the 
bulk of the nitrogen was contained in the austenite. 
The chemical analysis indicated that the nickel car¬ 
bide did not constitute the bulk of the material in 
this case as did the iron carbide (cementite) in the 
case of the white layers in other guns. In this 76-mm 
gun the surface layer was only slightly enriched in 
nickel, for the ratio of nickel to iron was not much 
greater for the surface layer than for the steel. 

The incipient melting points of some of the prod¬ 
ucts in this 76-mm gun tube were determined by the 
method described in Section 5.4.3. There is no doubt 
that liquefaction of the bore surface took place in this 
tube and therefore the temperatures at the bore sur¬ 
face during firing must have been above the incip¬ 
ient fusion temperatures of the products. For this 
reason these melting point determinations are of 
interest. 105 

The segregated surface layer flakes began to melt 
at 1150 C. This temperature may be somewhat higher 
than the actual incipient melting point of the white 
layer for minor amounts of material may have been 
dissolved in segregating the layer. Another experi¬ 
ment was carried out which was not subject to this 
error. Some bore surface material was detached 
mechanically from the area which had a thick white 
layer. This began to melt at 1125 C thus indicating 
that fusion of the reaction products could take place 
at a temperature at least 300 degrees lower than that 
of gun steel. 


12 6 SUMMARY OF EVIDENCE OF THERMAL 
AND CHEMICAL ALTERATION 

Generalizations concerning the mode of formation 
of the altered surface layers in eroded gun bores can 
be drawn from the nature of the erosion products 
described in the preceding sections of this chapter. 
The causes of gun erosion, which are discussed in 
Chapter 13, must take into account a mechanism 
which is in accord with the observed changes that 
take place in the bore surface of a gun during firing. 
The accumulated evidence shows that thermal, chem¬ 
ical, and mechanical factors have all played a role in 
the alteration of the bore surface. The first two of 
these are mostly important in producing the types of 
products that are found. The third is mostly effective 
in removing them. 

The bore surface of a gun is influenced to a greater 
depth by the heat developed in firing than it is by the 


chemical action of the powder gases. All but the sur¬ 
face of the altered layer has been shown to differ from 
the unaltered steel only in texture, not in chemical 
composition. It is believed, however, that at the very 
surface the altered layer has dissolved sufficient car¬ 
bon or nitrogen or both to liquefy partially and to 
retain austenite on quenching. Thus the first step in 
the chemical alteration of the bore surface is probably 
the formation of a surface layer which has the charac¬ 
teristics of the inner white layer when quenched. 

Once an inner white layer is formed, reactions at the 
bore surface are with this layer, which is partially 
liquefied in place during subsequent firing. With in¬ 
creasing chemical complexity of the immediate sur¬ 
face material, the melting temperature is decreased 
to the point where the material can flow. Material 
that is of a different chemical composition than the 
unaltered steel has been observed on many eroded 
bore surfaces. This forms what is known as the outer 
white layer. 

The surfaces of guns having such a layer show fea¬ 
tures characteristic of liquefaction, as described and 
illustrated in Section 10.5.2. Metallographic examin¬ 
ation of cross sections also gives evidence of liquefac¬ 
tion by revealing laminated structures, as if formed 
by superposed flows. These structures have been ob¬ 
served for both types of outer white layer. The one 
that consists essentially of cementite blades that ap¬ 
parently have crystallized out of a melt flows like a 
mush, whereas the rarer seemingly structureless one 
without cementite blades apparently flows more eas¬ 
ily. Thus all the accumulated evidence indicates that 
the outer white layer has been liquefied and has flowed 
into place. Its temperature of liquefaction for one gun 
has been found to be about 1125 C. 

That chemical reactions have taken place between 
the bore surface and the powder gases is borne out by 
x-ray studies made of the bore surface and of material 
found entrapped in the coppering. Further evidence 
for these reactions has been found by making detailed 
studies of the outer white layers. The difference in 
chemical composition between the outer white layer 
and the underlying steel is obvious from thermal con¬ 
siderations alone: thus the structure of this layer is 
not obviously changed by tempering; also partial 
melting of the material at a temperature much lower 
than that of gun steel as mentioned above indicates 
that the addition of materials to the surface has low¬ 
ered the fusion range. In the case of guns fired with 
single-base powders, the penetration of carbon and 
nitrogen has been proved, as discussed at length in 


CONFIDENTIAL 



SUMMARY OF THERMAL AND CHEMICAL ALTERATION 


259 


Chapters 13 and 14. Chemical analyses of outer white 
layers removed from bore surfaces by chemical means 
have confirmed the x-ray evidence that the bulk of 
the material removed in this manner was essentially 
a pure iron carbide associated with lesser amounts of 
iron nitrides. X-ray analysis of material removed by 
mechanical means showed that austenite, which usu¬ 
ally is dissolved when chemical methods of segrega¬ 
tion are employed, is an important constituent of the 
outer white layer. 


Guns fired with double-base powders rarely have 
outer white layers that are as readily amenable to 
detailed study as those in guns which have been fired 
with single-base powders. In this case, nevertheless, 
x-ray analysis of the bore surfaces and of decoppering 
residues, have shown that ferrous oxide and austenite 
are the dominant products. Thus reactions do take 
place between the bore surface and the gases from 
powders containing nitroglycerin as well as those con¬ 
taining only nitrocellulose. 




CONFIDENTIAL 



Chapter 13 

THE CAUSES OF GUN EROSION 

By C. A. Marsh a and J. N. Hobstetter b 


131 INTRODUCTION 

131,1 Experimental Approach 

T heories concerning the causes of gun erosion were 
formed as early as fifty years ago. 0 Since then 
many investigators using different techniques have 
arrived at different conclusions. The lack of agree¬ 
ment may be attributed to several reasons, the most 
important of which is the fact that most of the early 
contributors to this study tried to single out the dom¬ 
inant mechanism of erosion and to base their theories 
of gun erosion upon it alone. 

Gun erosion is a complicated process which in¬ 
volves a number of interrelated factors. In spite of 
their interdependence, which will be stressed in this 
chapter, different ones of these factors may predom¬ 
inate even in the same gun at certain positions in the 
bore. Since the early investigators studied several 
types of guns all with different firing histories, it is 
not surprising that the dominant mechanisms postu¬ 
lated did not always agree. 

Some of the disagreement among the early theories 
was undoubtedly due both to the techniques em¬ 
ployed, some of which are mentioned in Section 
12.1.1, and to the fact that in many cases only one 
line of experimentation was carried out. A variety of 
methods should be used to study a problem as com¬ 
plicated as gun erosion; furthermore, as many of 
these as possible should be resorted to in the case of 
one particular gun. Only after a number of guns have 
been studied systematically, is it possible to obtain a 
logical explanation of gun erosion. 

Such studies of eroded guns by Division 1 contrac¬ 
tors have been described in Chapters 10 and 12. 
Other pertinent information has been derived from 

a Geophysical Laboratory, Carnegie Institution of Washing¬ 
ton. (Present address: U. S. Geological Survey, Washington, 
D. C.) 

b Department of Metallurgy, Harvard University. (Present 
address: Department of Engineering Sciences and Applied 
Physics, Harvard University.) 

c The subject of this chapter is one on which extensive 
studies have been made. Since this was not intended to be a 
historical account, scarcely any reference has been made to 
work prior to 1940. Summaries of the early work may be found 
in three reports 16 - 261 * 479 that contain extensive bibliographies. 


examinations carried out concurrently at Watertown 
Arsenal on eroded gun tubes of different sizes, 282 - 283 
especially the new 76-mm gun, Ml. 247 - 248 - 257 - 259 - 208 In 
these examinations special attention 264-267 - 269 was 
paid to a type of failure designated as “progressive 
stress-damage” (Section 13.5.3). In addition, experi¬ 
ments on rifles, liners, vent plugs, test rods, and 
blocks of steel and other materials, described in 
Chapters 11, 14, 15, and 16, have yielded a great deal 
of information that can be fitted together to give a 
clear and consistent picture of the erosive process in 
guns. 

13.1.2 Principal Factors Involved 

It has proved of great value to group the many 
factors operating in erosion under three general head¬ 
ings: thermal, chemical, and mechanical factors. 
However, it is not to be assumed that these groups of 
factors act independently. Rather, it appears that 
gun erosion is caused by the simultaneous interaction 
of thermal, chemical and mechanical influences. 

Thermal factors, which come about by the transfer 
of heat from the hot powder gases to the bore walls, 
are found to contribute to erosion in three important 
ways. First, the heating causes softening of the bore 
which makes it more susceptible to the action of 
mechanical factors. Second, the heating causes the 
formation at the bore surface of a layer of austenite, 
which appears to have considerably less resistance to 
chemical attack by the powder gases than other steel 
modifications. Third, the heating causes a liquefac¬ 
tion at the bore surface of steel itself, when powders 
of high flame temperature are fired, or of low-melting 
mixtures of reaction products, when powders of low 
flame temperature are fired. 

Chemical factors, which come about through chem¬ 
ical interactions of the bore surface and constituents 
of the powder gases, are controlled as to nature and 
rate by the temperature obtaining. It is found that 
the highest range of temperatures favors oxidation, 
with the formation of FeO, while somewhat lower 
temperatures favor carburization, with the formation 
of Fe 3 C. Carbon and nitrogen are both found to pene¬ 
trate the outer skin of the austenitic layer, which is 


260 


CONFIDENTIAL 



THERMAL FACTORS 


261 


partly stabilized thereby. Continued reactions and 
liquefaction of this layer may result in the formation 
of a complex “white layer” on the bore surface. 

Mechanical factors, which involve the stressing of 
the bore surface by the pressure of the propellant 
gases and the projectile, effect the actual removal of 
solid and liquefied material from the bore. Sweeping 
and scouring action by the gases are found to be im¬ 
portant mechanisms. Scoring by gas leakage assumes 
importance in some cases. Abrasion and swaging by 
the projectiles are found to contribute appreciably to 
erosion during whatever time interval the bore ex¬ 
periences softening. 

Cracking of the bore is also a factor in causing gun 
erosion; and the same factors that cause erosion con¬ 
tribute to cracking. Since it is important both as 
cause and effect, it is treated in a separate section, 
where its apparent dependence upon very numerous 
and intricate interrelations among the influences of 
thermal, chemical, and mechanical shock is de¬ 
scribed. 

The term “shock” has been applied to the chemical 
as well as to the thermal and mechanical factors to 
emphasize the briefness of the interval of 'time in 
which all three of these factors usually act to alter the 
bore surface in such a manner that it may be readily 
removed or distorted. 

132 THERMAL FACTORS 

13 21 General Statement 

Under the heading of thermal factors are grouped 
all of the effects directly associated with either the 
attaining of elevated temperature levels within the 
bore walls or the changing of those levels during the 
firing of a gun. Evidently, these effects contribute 
directly to gun erosion only insofar as they include 
liquefaction of material at the bore surface. Never¬ 
theless thermal factors play a most important role in 
erosion when they act in concert with chemical and 
mechanical factors. Indeed, they may be said to act 
as overall regulators of the erosive process. 

For example, the effect of mechanical stresses de¬ 
pends upon the mechanical properties of the bore wall 
at the moment when it experiences the stresses. These 
properties depend in turn on the temperature distri¬ 
bution in the barrel at that moment. Again, the effect 
of chemical reaction of the bore surface depends on 
the nature and rate of reactions between the bore 
material and the constituents of the powder gas. 


Which of the many possible reactions actually occur 
at any moment, as well as the rate at which they pro¬ 
ceed, depends upon the temperature obtaining at the 
bore surface at that moment. 

It is thus evident that any adequate understanding 
of the cause of gun erosion must include as a first step 
a full history of temperature changes experienced by 
the bore wall. In other words, thermal factors must 
be quantitatively understood. 

13.2.2 Thermal Softening of the Bore 

It is well known that the hardness of gun steel 
drops rapidly if its temperature is raised much above 
500 C and that concurrently the strength decreases 
and the ductility increases. If stresses are applied to 
the steel when it is in this softened condition, it is 
obvious that considerably different behavior is to be 
expected than if they were applied to cold steel. It is 
important to consider, then, what kinds of stress the 
bore of a gun experiences when its surface is thermally 
softened. Two extreme cases can be pictured which 
bracket the true behavior of any gun: single-shot 
guns in which cooling is substantially complete be¬ 
tween rounds and rapid-fire guns in which the whole 
bore wall may rise above the softening temperature 
during prolonged bursts. The first few shots fired in 
rapid succession from an initially cold machine gun 
barrel represent an intermediate case: here the whole 
thickness of the barrel has not yet become hot, but 
near the origin of rifling a very thin layer of steel be¬ 
ginning at the bore surface remains so hot between 
rounds that it is soft, and variously transformed on 
cooling. 

Single-Shot Guns 

It is not difficult to imagine the history of a single 
round from among many fired in a single-shot gun. 
At the moment of firing the bullet begins to move 
down the bore, passing over a surface bearing the 
hardened layer which is discussed in Section 13.2.3. 
This hard layer probably resists adequately the stres¬ 
ses accompanying the engraving and friction of the 
bullet except that it may become cracked because of 
the mechanical shock and its low ductility. Such 
cracking is discussed in detail in Section 13.5.3. 

Hot gases follow immediately behind the bullet 
and heat the bore surface. It has been shown in Chap¬ 
ter 5 that a relatively thin skin at the bore surface is 
rapidly heated in this way to temperatures far be- 


CONFIDENTIAL 



262 


THE CAUSES OF GUN EROSION 


yond the softening temperature. The actual amount 
of softening that results from this heating increases 
progressively, since softening is a function of the time 
at temperature. In any case, as the bore softens it is 
subjected to scouring action and gas pressure, as de¬ 
scribed in Section 13.4.1, and is more easily affected 
by these processes because of deteriorating mechani¬ 
cal properties. 

It may be noted that some gases may leak past the 
bullet and preheat the bore. Friction between the 
bullet and the bore also cause heating of the bore. 
Whatever softening accompanies these phenomena, 
of course, permits an increased superficial deforma¬ 
tion of the bore surface during engraving and passage 
of the bullet, but it seems unlikely that the effect 
could be large except in extreme cases of faulty 
obturation. 

Finally, the bullet and gases rush from the gun, the 
heat transferred to the bore surface concurrently dif¬ 
fuses very rapidly into the cold metal of the bore wall, 
and the hardness of the bore surface is restored. 

Rapid-Fire Guns 

In principle, the effects accompanying rapid fire 
are the same as those accompanying single shots ex¬ 
cept that the heat transfer to the bore surface is ef¬ 
fected so frequently, as described in Sections 5.4.2 
and 5.5.1, that the whole barrel begins to heat. As is 
brought out under “Swaging” in Section 13.4.2, a layer 
extending to a considerable depth below the bore sur¬ 
face may remain in the soft austenitic state. Under 
these conditions, the bullet passing down the bore 
encounters only a very soft surface. In addition, the 
gas pressure may actually dilate the whole softened 
tube. 


d This section is based on an NDRC report 124 by one of the 
authors of this Chapter. It uses evidence obtained from a study 
of eroded gun barrels and liners fired in a particular gun under 
hypervelocity conditions to support a theory of the interrela¬ 
tions between thermal and chemical factors in the erosion of 
that gun. There has not been opportunity to determine the 
extent to which the same approach may be made in the study 
of other eroded guns. Since this is the first attempt to reduce 
this important subject to a quantitative basis, it seems desir¬ 
able to present this point of view in some detail, even though 
full experimental verification is lacking. Thus the attempted 
experimental comparison of the relative reactivities of austen¬ 
ite and ferrite is open to the criticism that the observed differ¬ 
ences in chemical reactivity can be accounted for on the 
grounds of differences in temperature and texture. (See Section 
12.1.2.) Also examination 86 of some medium-caliber guns fired 
at conventional velocities has revealed “mushrooming” below 
chromium plate on a steel surface that had not been thermally 


13.2.3 Thermal Transformation at the 
Bore Surface d 

Development of Austenitization 

The methods of calculation outlined in Section 
5.4.1 show that the temperature of the bore surface of 
gun tubes can rise not only above the softening tem¬ 
perature, but above the critical temperature of steel, 
about 750 C. Indeed, it has been shown that at the 
hotter parts of the bore surface, a zone more than 
0.002-in. thick may experience supercritical temper¬ 
atures during the firing of a single round. This cir¬ 
cumstance suggests very strongly that the bore sur¬ 
face can be transformed to the high-temperature 
(gamma) modification of iron known as austenite. 

That such transformation does actually occur is 
abundantly demonstrated by the existence of the 
thermally altered layer which is found in all gun steel 
tubes, as has been described in Section 12.1.2. It 
may be mentioned again here that in the caliber .50 
erosion-testing gun (Section 11.2.1) this layer is 
formed by the firing of one round, that its thickness 
depends almost entirely on the powder and the con¬ 
ditions of firing, as is emphasized in Section 15.5.4, 
and that it forms under thin protective coatings 
which prevent direct contact with the powder gases. 
These facts argue convincingly that the layer is of 
thermal, not chemical, origin. 

Mechanism. Any thermal transformation of this 
sort depends on time as well as temperature. The 
transformation in gun steel can be pictured as involv¬ 
ing first the formation of austenite from the ferrite 
(alpha iron modification containing little carbon) and 
the subsequent solution of carbides in austenite as 
solvent. Both processes are controlled by the diffu¬ 
sion rate and take time to occur. Of course, it would 
be proper to speak of austenitization even if the car¬ 
bide solution process were incomplete, and thermally 
transformed layers containing undissolved carbides 
have been observed after the firing of only one round 
in the caliber .50 erosion-testing gun (Section 11.2.1). 

Evidently, the region along the bore surface that is 
transformed is the region which experiences thermal 
conditions sufficient to cause austenitization and the 
rather sharp interface between altered and unaltered 
steel experiences critical conditions. Because this in- 


altered. It is hoped, at any rate, that future investigators may 
be stimulated by this account of a new approach to the subject 
to pursue it further. (Editor’s note.) 


CONFIDENTIAL 






THERMAL FACTORS 


263 


terface reaches supercritical temperatures during one 
round, time at temperature must be the limiting 
factor. 

A study was made of the time interval during 
which the altered layer interface was at supercritical 
temperatures during single rounds fired with different 
powders and charges in the erosion-testing gun, 
which is described in Section 11.2.1. Although the 
austenitization time could not be definitely evaluated, 
it seemed to be constant. It was determined, however, 
that the thermally altered layer was formed in all 
regions that reached the critical temperature at least 
0.0002 sec before the attainment of maximum tem¬ 
perature at the bore surface. This criterion enables 
the thickness of the altered layer in different guns to 
be predicted under a wide variety of firing conditions. 
It also implies that the austenitization time is of the 
order of 0.5 msec. 

Distribution. Thermally altered layers in the gun 
barrels studied were found to vary in thickness both 
circumferentially and axially, as described in Section 
15.3.3. Thicker layers were found on the protruding 
lands, particularly on the land corners where heat can 
enter the metal through two adjacent surfaces. Im¬ 
poverishment of the heat transfer to the grooves 
immediately next to the lands was such that the over¬ 
all heat transfer was inappreciably different from 
what it would be in unrifled tubes. 48 

The thickness of the layer decreased toward the 
muzzle. Although the layer formed during the firing 
of the first round, it thickened progressively with con¬ 
tinued single shots except near the origin of rifling. 
Thus, nearly complete cooling between rounds must 
have been effected near the origin of rifling by the 
relatively thick, cold bore walls, whereas toward the 
muzzle incomplete cooling was effected by the thinner 
walls. 

Reactivity of Austenite 

The thermally altered layer in single-shot guns 
occurs in two modifications, both of which differ from 
the unaltered steel. The layer is austenite when hot, 
as described above, and martensite (together with 
some retained austenite) when it has cooled quickly 
after the firing of a round, as described below. Either 
or both of these modifications may be expected to 
have different chemical properties from the ferrite of 
the unaltered steel and it is important to consider 
these properties since, in any round after the first, it is 
the altered layer that is eroded by the powder gases. 


Observation of Greater Reactivity. One of the earliest 
observations 86 - 112 relative to the behavior of barrels 
plated with erosion-resistant coatings was that if the 
plating were insufficiently thick, it was undercut by 
considerable erosion of the thermally altered layer 
that formed beneath it. This behavior is shown quite 
clearly in Figure 1. The thin coating is chromium 
plate which has become cracked during firing. The 
cracks have provided the powder gases access to the 
thermally altered layer which has, thereupon, devel¬ 
oped a pocket-like erosion called “mushrooming.” 



Figure 1 . Pocket-type erosion beneath chromium 
plate of a steel liner fired in the erosion-testing gun. 
Nital etch. 200X. (This figure has appeared as Figure 
10 in NDRC Report No. A-452.) 


In no case has it been observed that “mushroom¬ 
ing” penetrates the unaltered steel. The increased 
temperature near the bore surface is insufficient to 
explain either this greater reactivity or the rather 
sharp drop in reactivity at the altered layer interface. 
Accordingly, it may be concluded that the thermally 
altered layer is in fact much more easily eroded by 
the powder gases than is unaltered steel. 

It remains to show if either or both modifications 
of the altered layer are easily eroded. The reactivity 
of the martensitic layer was studied by forming first 
an altered layer under a thin resistant coating, then 
protecting this layer with a coating more than crit¬ 
ically thick which permitted no further austenitiza¬ 
tion of underlying steel. (Studies to determine the 
critical thickness of chromium-plate are described in 
Section 31.5.) The results are shown in Figure 2 
where no evidence of “mushrooming” in the ther¬ 
mally altered layer can be seen. Thus it follows that 


CONFIDENTIAL 






264 


THE CAUSES OF GUN EROSION 


only in its austenitic state is the altered layer much 
more easily eroded than ferrite. 

Relation to Chemical Erosion. On the basis of these 
findings it is not difficult to draw a picture of the 
process of chemical erosion insofar as it is related to 
thermal factors. Let us consider any one round fired 
in a single-shot gun. The erosive gases will first strike 
the bore when it is rather cold and only reactions that 
can take place at low temperatures may be expected 
to occur. The bore surface will be heated rapidly, 
however, first softening and then being transformed 



; i 



Figure 2. Chromium-plated liner fired in the erosion¬ 
testing gun. The altered layer was first formed under 
a thin plate and subsequently protected by the deposi¬ 
tion of a very thick plate. Nital etch. 200X. (This figure 
has appeared as Figure 11 in NDRC Report No. A-452.) 

to austenite. It is believed that during and after this 
transformation the rates of whatever chemical reac¬ 
tions are going on are enormously accelerated, on one 
hand by the higher temperature but mostly by the 
formation of the more reactive phase. 

Eventually, the bore-surface temperature reaches 
its maximum and falls back below the critical tem¬ 
perature whereupon most of the austenite decom¬ 
poses. The erosive reactions, in turn, become much 
less violent along with the disappearance of austenite 
and gradually change character as they revert to the 
low-temperature type. 

The nature of all these reactions is discussed in 
detail in Section 13.3 on “Chemical Factors.” 


Decomposition of Austenite 

The decomposition of austenite which occurs after 
the peak temperature at the bore surface has been 
reached may be expected to take place in the same 
manner as in the ordinary heat treatment of steel. 
All of the structures that may be obtained by differ¬ 
ent heat treatments of steel are shown diagrammati- 
cally in Figure 3. The mathematical study of wall 
temperatures given in Section 5.4.1 reveals that the 
temperature of the bore surface in single-shot guns falls 
very rapidly indeed. All such very rapid cooling rates 
suppress completely the isothermal decomposition 
of the austenitic steel of the composition of gun steels 
so that pearlites do not form. Instead, the martensite 
reaction sets in at a very low temperature, which is 
independent of the cooling rate and the hard, marten¬ 
sitic phase is formed. This phase, then, is what is 
actually observed when the altered layer is studied 
metallographically. 

It may be noted that the temperature of the begin- 
ing of the martensite reaction is generally depressed 
whenever foreign elements are dissolved in the aus¬ 
tenite. Carbon and nitrogen have a pronounced effect 
in this regard and their presence in the austenite can 
easily depress the temperature of the martensite re¬ 
action below room temperature. 

In Chapter 14 it is shown that carbon and nitrogen 
do penetrate slightly the austenitic layer. In this 
manner, the immediate surface of the austenite usu¬ 
ally becomes stabilized and does not form martensite. 
This thin skin is usually called the inner “white” layer. 
The outer “white” layer which is sometimes abun¬ 
dantly found also contains austenite. For a detailed 
description of the “white” layers see Section 12.1.2 . 

13 2 4 Liquefaction of the Bore Surface 

Evidence of liquefaction at the bore surface has 
been presented in Sections 10.5.2 and 12.6. Such a 
process, assisted by motion imparted mechanically, 
is the most direct way in which erosion can be brought 
about by thermal factors. Even so, liquefaction is 
aided to a very great extent if low-melting mixtures 
of chemical reaction products are first formed at the 
bore surface.® The nature and the rate of the liquefy¬ 
ing process are most easily classified according to the 
thermal conditions obtaining in the gun and these 


e An experiment to evaluate the relative importance of 
melting and chemical attack in the erosion of vent plugs is 
described in Section 16.4.14. 


CONFIDENTIAL 






THERMAL FACTORS 


265 



Figure 3. The microstructural constituents of steel. (This figure has appeared as Table XX in NDRC Report No. A-91.) 


CONFIDENTIAL 




























266 


THE CAUSES OF GUN EROSION 


conditions are largely determined by the type of pow¬ 
der fired. The two large classes that can be distin¬ 
guished are discussed in the following paragraphs. 

Liquefaction with Double-Base Powders 

It has been found in experiments with the caliber 
.50 erosion-testing gun (Section 11.2.1) that measur¬ 
able erosion of gun steel barrels begins with the first 
round when double-base powders of high flame tem¬ 
perature are fired. 123 Metallographic examination has 
also shown evidence of liquefaction after the firing of 
the first round. 123 Furthermore, calculations show 
that the peak temperature of the bore surface is well 
above the melting point of steel when such powders 
are fired. 124 The conclusion that liquefaction of an 
ordinary gun steel surface takes place when very hot 
powders are fired is, therefore, inescapable. Any low- 
melting mixtures of reaction products which may also 
be formed at the bore surface will, of course, aid the 
liquefaction, but it would appear that these mixtures 
are not necessary and that liquefaction can occur in 
their absence. 

It follows, then, that the rate of liquefaction should 
be limited by the rate at which heat can be trans¬ 
ferred to the crystalline part of the bore surface and 
not by the rate at which chemical reaction products 
are formed and removed. Erosion may be expected to 
occur at a rapid rate starting with the first round, 
which expectation is in accord with observed facts. 

Liquefaction with Single-Base Powders 

When single-base powders of low flame tempera¬ 
ture are fired in the caliber .50 erosion-testing gun, 
the observed facts of erosion are quite different. 
Measurable erosion does not begin with the first 
round, rather there is a lag or “incubation” period. 123 


Metallographic evidence of liquefaction is not found 
until the end of the “incubation” period when peb¬ 
bling of the surface (Section 10.5.2) begins. 123 Calcu¬ 
lations show that the peak bore-surface temperature 
is less than the melting point of steel. 124 These facts 
can be explained only on the basis that a mixture of 
chemical reaction products and not gun steel itself is 
liquefied. The “incubation” period then becomes a 
period during which initial reaction products are be¬ 
ing formed and the subsequent erosion is limited by 
the rate of continuing formation and removal of these 
mixtures and not by the relatively rapid rate of heat 
transfer that results in their liquefaction. 

13 2 5 Importance of Thermal Factors in 
the Centering Cylinder 

Considerable erosion often takes place in the cen¬ 
tering cylinders of guns. Here the main cause of ero¬ 
sion is the heat transferred from the powder gases, 
since the engraving of the projectile has not yet be¬ 
gun. Because of the contour of the chamber, the flow 
of gases in the centering cylinder is turbulent. Thus 
more heat is imparted to the walls here than in the 
bore where the flow of gases, next to the walls at 
least, is laminar. The type of flow in the different 
portions of a gun tube is illustrated in Figure 4. The 
gas stream converges in a vena contrada effect when 
it passes from the large portion of the chamber, and 
the full force of the turbulent stream thus produced 
does not strike the tube walls until it reaches an area 
within the centering cylinder. Here the turbulent 
gases strike the surface in a radial normal direction 
and eddy backwards. 

Although erosion is produced by these turbulent 
gases in both bag-fired and case-fired guns, its effects 
are more noticeable in the latter because erosion of 
the chamber does not occur in the area protected by 



Figure 4. The transfer of heat from the turbulent powder gases to the bore surface is increased by the vena contractu 
effect introduced by chambrage. 


CONFIDENTIAL 




















CHEMICAL FACTORS 


267 


the cartridge case. In extreme cases, the centering 
cylinder, between the origin of rifling and the position 
of the cartridge case, appears to have been scooped 
out. 

13 3 CHEMICAL FACTORS 

1331 Introduction 

The discussion of the products of erosion in Chap¬ 
ter 12 has shown that a consideration of the reactions 
of the constituents of the powder gases with gun steel 
is of utmost importance in formulating a mechanism 
of gun erosion. The elements present in the gases re¬ 
sulting from the combustion of all the usual propel¬ 
lants are carbon, oxygen, hydrogen, nitrogen, and 
sulfur. The effects of separate ones of these constitu¬ 
ents has been studied by means of carefully controlled 
laboratory experiments, the results of which are given 
in Chapter 14. 

It must not be concluded, however, that these ele¬ 
ments act independently as such to bring about 
chemical changes in the bore surfaces of guns during 
firing. Rather, the powder gases must be considered 
to consist of a mixture of molecules in which the 
above elements are combined. Their concentrations 
depend not only on the type of powder but to a large 
extent on the temperature of the gas mixture. The 
type and rate of reactions which take place between 
the steel and the powder gas depend essentially on 
the temperature of the bore surface and the compo¬ 
sition of the gas, which in turn depends on its temper¬ 
ature. These reactions and the conditions effecting 
arid affecting them are the subject of this section. 
Insofar as possible they are related to the observa¬ 
tions and results described in Chapters 12 and 14. 

13 3 2 Penetration of Carbon and Nitrogen 

Difference Between Penetration and Surface 
Reactions 

The reactions which produce carbides and nitrides 
are considered to be the most important of those 
which contribute to gun erosion. This conclusion was 
derived not only from the evidence presented in 
Chapter 12 but from thermodynamic considera¬ 
tions 60 which showed that in guns fired with single¬ 
base powders the major product to be expected is 
cementite (Fe 3 C), whereas in guns fired with double¬ 
base powders the major product is wiistite (FeO). In 


the latter case, however, the oxidation of the bore 
surface is probably of little importance since heat is 
sufficient to melt the stee} under an oxide film, as 
was mentioned in Section 13.2.4. Furthermore car¬ 
burization and nitriding of the steel are penetration 
reactions, whereas oxidation is essentially a surface 
reaction. Oxide films are easily removed in firing, but 
the metallic alteration products containing carbides 
and nitrides may accumulate in some cases to a con¬ 
siderable thickness; hence, in the case of oxidation 
evidence is not usually as abundant as in the case of 
penetration of carbon and nitrogen. 

These penetrating reactions produce complications 
in the determination of ultimate products by ther¬ 
modynamic methods, as is described later, for solid 
solutions involving carbon and nitrogen are formed, 
for which there are insufficient data to permit of their 
treatment. It is believed that an understanding of the 
systems Fe-C and Fe-N so far as is known is desirable 
in order to consider the role of carbon and nitrogen in 
gun erosion. 

Iron-Carbon System 

A simplified, equilibrium diagram 503 for the iron- 
carbon system is shown in Figure 5. This might be 
more truly called a diagram for the iron-iron carbide 
system since Qnly the high iron end is considered. 



Figure 5. Iron-carbon constitution diagram. (Repro¬ 
duced from Metals Handbook, 1939 ed., p 368, by per¬ 
mission.) 


CONFIDENTIAL 










































268 


THE CAUSES OF GUN EROSION 


This diagram is simplified in that it gives no indica¬ 
tions of the microstructural constituents, which are 
combinations of the crystalline phases arranged in 
definite textural patterns that are observed metallo- 
graphically. For example, the lower half of the dia¬ 
gram shows only the phases ferrite and cementite 
(6.67% C). In the field designated as ‘‘ferrite plus 
cementite” the microstructural constituents pearlite 
and ferrite are present to the left of a perpendicular 
dropped from S ; between perpendiculars from S and 
E, pearlite and cementite are present. Pearlite itself 
consists of a mixture of ferrite and cementite but in 
definite proportions and with a characteristic struc¬ 
ture. 

For those who are not familiar with such constitu¬ 
tion diagrams several examples will be given to illus¬ 
trate the meanings of the various points, lines, and 
fields. The left edge of the diagram represents pure 
iron. As the right side of the diagram is approached 
the composition is that of iron with increasing 
amounts of carbon. The line A BCD, called the li- 
quidus, represents the beginning of solidification on 
cooling and the end of melting on heating. All points 
above this line represent alloys in a completely mol¬ 
ten condition. All points below ABCD represent 
alloys partially or completely solid. 

The alloy containing 4.3% carbon is the eutectic 
alloy and solidifies entirely at the point C with the 
simultaneous formation of austenite and cementite. 
At the eutectic temperature austenite will hold 1.7% 
carbon. Cementite precipitates out on cooling below 
this temperature, the limit of solubility of cementite 
in austenite being represented by the line SE. The 
point S is called the eutectoid. It is similar to the eu¬ 
tectic C except that it is completely surrounded by 
solid fields. 

Any further explanation of this diagram wall be 
postponed until it is referred to in connection with 
the phenomena associated with gun erosion. Infor¬ 
mation on microstructures obtained by quenching 
and by other methods which depart from equilibrium 
conditions may be obtained from Figure 3. 

Iron-Nitrogen System 

Several different diagrams for the system iron- 
nitrogen or iron-iron nitride have been proposed. 503 
Molecular nitrogen does not react appreciably with 
iron; the formation of nitrides is usually carried out 
through the medium of ammonia. The constitution 
diagrams do not represent equilibrium conditions but 


must be considered behavior diagrams. The work on 
this system is far from complete, but enough has been 
done to show the similarity between alloys of iron 
and carbon and alloys of iron and nitrogen. 40 Thus 
the names of the following constituents that have 
been found are self-explanatory: nitroferrite, nitro- 
austenite, and nitromartensite. The eutectoid alloy 
has been called braunite. 

Some of the proposed diagrams were based on ex¬ 
perimental work which employed the x-ray diffrac¬ 
tion method of analysis. The pertinent data obtained 
are given here. The eutectoid temperature was found 
to be 591 C with a nitrogen content of 2.35%. The 
occurrence of iron nitrides, Fe 4 N and Fe 2 N^, has been 
established. The former, which is called the “gamma- 
prime phase,” has a face-centered cubic structure 
with a cell size somewhat larger than that of the 
“gamma phase” (nitroaustenite) when it contains al¬ 
most as much nitrogen. The latter, which is called the 
“epsilon phase,” has a hexagonal structure and may 
contain from 8 to 11.3 per cent of nitrogen. The com¬ 
pound Fe 4 N is supposed not to be stable at high 
temperature and to decompose at 650 C into the 
gamma and epsilon phases. 

Further experimental work was carried out on the 
system iron-nitrogen in connection with the investi¬ 
gation of gun erosion by Division 1. This work 28 is 
discussed in Section 14.5.3. The experimental results 
led to the conclusion that since they are not stable 
under firing conditions, iron nitrides are formed in 
guns during the cooling stage either by low-tempera¬ 
ture reactions with the powder gases or by precipita¬ 
tion from nitroaustenite. 

Penetration of Carbon and Nitrogen 

If conditions are such during firing that carburizing 
reactions may take place, then carbon may penetrate 
the steel bore-surface. Since carbides and nitrides are 
generally found together in the eroded bore surface 
and since the two systems of those elements with iron 
are somewhat similar, it is assumed that nitrogen and 
carbon penetrate simultaneously. Confirmation in 
the case of carbon was obtained from the work dis¬ 
cussed in Section 14.2 in that carbon penetration was 
observed, except when FNH-M2 powder was fired, 
which involved oxidizing conditions. 

When austenitization of the bore takes place, as 
described in Section 13.2.3, it is to be expected that 
the surface may pick up carbon or nitrogen or both 
from the powder gases, since the steel does not con- 


CONFIDENTIAL 



CHEMICAL FACTORS 


269 


tain the total amount of carbon or nitrogen that may 
dissolve in austenite. That carbon actually does pen¬ 
etrate below the layer of reaction products was 
proved by the experiments with radioactive carbon, 
the results of which are given in Section 14.2. The 
carbon did not penetrate to the full depth of the ther¬ 
mally altered layer. 

The penetration of carbon or nitrogen or both [at 
least as far as the bottom of the inner white layer 
(described in Section 12.1.2) while it remained at the 
bore surface] is inferred from the fact that this layer 
is more resistant to mild etches than is the thermally 
altered layer beneath it. Also, it appears to have been 
locally slightly liquefied in place, which signifies that 
the melting point has been lowered by additions of 
material from the powder gases. The fact that an 
apparently austenitic structure has sometimes been 
observed is considered as evidence that austenite was 
retained in this layer because it had dissolved carbon 
or nitrogen or both. 

The outer white layer, described in Section 12.1.2, 
although it consists essentially of cementite, is not 
believed to be merely the result of surface reactions 
but to constitute evidence of both carbon and nitro¬ 
gen penetration. It presumably was formed by con¬ 
tinued penetration of the inner white layer by the 
carbon and nitrogen of the powder gases, as described 
below. 

Formation of the White Layers Enriched in 
Carbon and Nitrogen 

Experiments showed in separate instances that 
both thermal 123 and a minor amount of chemical" 
alteration (carbon penetration) take place during the 
firing of the first round through a gun. f The amount 
of carbon which has penetrated increases with the 
number of rounds. By the time the surface has taken 
up enough carbon to be partially liquefied (see Figure 
5 for the boundaries of the field “austenite plus li¬ 
quid”)? there is no longer the question of whether 
compounds of iron or solid solutions are formed; the 
powder gases react with the partially liquefied surface 
which presumably can dissolve more carbon, and pos¬ 
sibly nitrogen, than austenite can. The result is an 
increased “penetration.” 


f It is not known whether this is true for all types of 
guns. The study of thermal alteration was carried out with 
the caliber .50 erosion-testing gun; the study of carbon pene¬ 
tration with caliber .30 rifles preheated to different tem¬ 
peratures. 


This process continues with a further number of 
rounds and the melting point is successively lowered 
until it may reach the eutectip temperature (point C 
in Figure 5). With the lowering of the melting point, 
the surface layer may actually flow. The evidence 
presented in Section 12.6 indicates that the outer 
white layer is material that has flowed. 

It appears in the case of guns tha # t have a white 
layer in which cementite is the primary constituent 
that the eutectic composition of carbon was exceeded 
and that cementite plus liquid (Figure 5) had been 
present on the surface. Nitrides are found associated 
with cementite in this layer which means that, what¬ 
ever the mechanism by which nitrogen was intro¬ 
duced, it did penetrate and nitrides crystallized either 
out of the melt or out of the austenite which was 
retained in the interstices between the cementite 
blades. 

For a further discussion of the role of nitrogen in 
gun erosion see Section 14.5. 


13 3 3 Reactions at the Bore Surface 
During Firing 

In the preceding section the systems iron-carbon 
and iron-nitrogen were discussed to orient the reader 
with respect to penetration studies. In the reaction of 
the gun steel with all the constituents of the powder 
gases, the system involved is Fe-C-O-H-N-S. All the 
details of this complicated multicomponent system 
are not necessary in determining what products are 
likely to be formed on bore surfaces during firing. 
The thermodynamic methods used to determine the 
theoretical products of erosion were described briefly 
in Section 12.2.4. The reactions involved are the 
subject of this section. 

Reactions with the Components of the Water 
Gas Reaction 

The compounds of iron that may be formed on the 
bore surface under a variety of conditions depend on 
concentrations of the components of the powder gases 
under those conditions. At the temperatures obtained 
during firing, the major components are those of the 
water gas reaction (1), 107 as is brought out in Section 
2 . 2 . 2 . 

C0 2 + H 2 ^CO + H 2 0 (1) 

The equilibrium constant for this reaction, expressed 


CONFIDENTIAL 




270 


THE CAUSES OF GUN EROSION 


by equation (2) [which is the same as equation (5) of 
Chapter 2], 

_ [CO] [ H 2 0] 

- [C0 2 ] [H 2 ] ’ y ’ 

and the C0/C0 2 ratio decrease with the temperature 
of the gases. At lower temperatures, soot formation 
[equation (3)] , 

2CO^C + C0 2 (3) 

is a competing reaction. The compositions of single¬ 
base and double-base propellants, discussed in Sec¬ 
tion 2.3.4, are such that the flame temperatures of the 
latter are higher and, for a given value of K wg , the 


The calculations were made for various gas densi¬ 
ties, bore-surface temperatures and pseudo-gas tem¬ 
peratures, which are those from which the gas was 
considered quenched. The dependence of the ultimate 
product on the variables just mentioned was repre¬ 
sented graphically by the partition of a three-dimen¬ 
sional space diagram. Figures 6 and 7 represent 
planes cut through such a figure at a certain gas 
density for single-base and double-base powder re¬ 
spectively. 

These graphs show that at the instant of firing the 
products obtained with single-base and double-base 
powders are Fe 3 C and FeO respectively. When the 
gases are not quenched, oxides (fields immediately to 



Figure 6. The ultimate product for FNH-M1 powder 
as a function of bore-surface temperature and pseudo¬ 
temperature of the gas at a density of 0.2 g per cu cm. 
(This figure has appeared as Figure IB in NDRC Re¬ 
port No. A-301.) 



900 1000 1100 1200 1300 1400 1500 1600 

PSEUDO-TEMPERATURE OF GAS IN DEGREES K 

Figure 7. The ultimate product of FNH-M2 powder 
as a function of bore-surface temperature and pseudo¬ 
temperature of the gas at a density of 0.2 g per cu cm. 
(This figure has appeared as Figure 3B in NDRC Re¬ 
port No. A-301.) 


C0/C0 2 ratios are lower. These relationships are rep¬ 
resented graphically in Figure 9 of Chapter 2. 

Standard thermodynamic methods were used to 
determine the ultimate products that would be 
formed if equilibrium were established between a 
finite quantity of iron and an infinite quantity of the 
powder gas. 60 In a competing series of reactions, the 
ultimate product is the one that requires the greatest 
negative free energy change for formation directly 
from iron. The compositions of the gas had to be cal¬ 
culated both for the temperature from which it was 
considered to be quenched and for temperatures ap¬ 
proaching those of the bore surface when it wa? con¬ 
sidered to have been cooled slowly so that equilibrium 
in the gas phase was established at those lower tem¬ 
peratures. 


right of line of unit slope) may be formed in the case 
of the single-base powders. In both cases Fe 3 0 4 is the 
oxide formed at the low surface temperatures ob¬ 
tained when the bore has cooled somewhat after fir¬ 
ing. Thin films of magnetite (Fe 3 0 4 ) have been found 
by electron diffraction analysis on the bore surfaces 
of guns fired with both types of powders. 137 The fact 
that magnetite was the dominant erosion product 
found in the cracks in an eroded gun was attributed 
to the steep thermal gradient between the bore and 
outside of the gun. 98 

It can be seen from the above results and from a 
consideration of the water gas reaction (1) that at the 
instant of firing, when the gases are quenched so that 
the equilibrium constant K wa is high, the reactions 
that take place in the case of double-base and of 


CONFIDENTIAL 































CHEMICAL FACTORS 


271 


single-base powders are probably those represented 
by equations (4) and (5), respectively. 

Fe + C0 2 ^FeO + CO (4) 

3Fe + 2CO^Fe 3 C + C0 2 (5) 

Some of the oxide formation may also be attributed 
to reaction with H 2 0 by equation (6), 

Fe + H 2 O^FeO + H 2 , (6) 

since double-base powders, which have lower C0/C0 2 
ratios than single-base powders, have a correspond¬ 
ingly low H 2 /H 2 0 ratio, according to equation (2). 
The formation of Fe 3 0 4 takes place by reaction (7) 

3Fe + 4C0 2 ^Fe 3 0 4 + 4CO (7) 

because of the shift in the equilibrium of the water 
gas reaction with a resulting decrease in the C0/C0 2 
ratio. Fe 2 0 3 , considered a possibility according to 
equation (8), 

2Fe + 3C0 2 ^Fe 2 0 3 + 3CO (8) 

was not found to be an ultimate product under any 
condition. 

The formation of Fe 3 C could take place in two 
stages, first the decomposition of CO into C0 2 [equa¬ 
tion (3) ] and then the reaction (9) 

3Fe + C^±Fe 3 C (9) 

of carbon with iron. The formation of Fe 3 C directly 
from CO according to equation (5), however, is more 
likely, since Fe 3 C is stable with respect to Fe, CO, and 
C0 2 under conditions such that C is not stable with 
respect to CO and C0 2 . 60 

Kinetics, of course, are not considered in the de¬ 
termination of ultimate products. It is quite likely, 
however, that they have some importance in the 
chemistry of gun erosion. Unfortunately nothing is 
known of the rates of the competing reactions that 
may take place at the bore surface, but it is consid¬ 
ered safe to say that the ultimate products may not 
always be the major products and perhaps more than 
one product may result. A reaction that may be of 
great importance is represented by equation (10). 

3FeO + 5CO^Fe 3 C + 4C0 2 . (10) 

A plot of the equilibrium curves for the important 
equations is shown as Figure 8. Curve F represents 
equation (10). Below curve B, the formation of FeO 
from Fe, the free energy change for FeO is negative. 
Above curve E, the formation of Fe 3 C from Fe, the 
free energy change for Fe 3 C is negative. Both FeO 


and Fe 3 C are stable with respect to Fe in the region 
between these two curves, although not equally stable. 
The free energy changes fo^- both compounds are 
equal along curve F. In a system corresponding to a 
point above this curve, Fe 3 C will be the major prod¬ 
uct if equilibrium conditions are fulfilled, and below 
the curve, FeO will be the major product. 

The significance of curve A [equation (3) ] on this 
plot is to point out that conditions existing at the 
surface of the metal may be carburizing even below 
curve F since carbon is precipitated above curve A. 
The rate of diffusion of carbon into iron is higher than 
that of oxygen; the formation of carbide rather than 
oxide will certainly be favored at any appreciable 
distance below the metal surface. With an increase in 



Figure 8. Equilibrium curves for reactions involved 
in the chemical alteration of gun steel by powder gases. 
(This figure has appeared as Figure 5 in NDRC Report 
No. A-466.) 

pressure, the formation of Fe 3 C from C [equation (9)] 
is favored. It can be seen that reactions (3) and (5) 
also proceed to the right; therefore, no matter by 
what mechanism Fe 3 C is produced, an increase in 
pressure favors its formation rather than that of FeO 
for a given C0/C0 2 ratio, since the reaction (4) to 
form FeO is not dependent on pressure. 

Reactions with Hydrogen Sulfide 

The only compound of sulfur that is present to any 
appreciable extent in the powder gases is hydrogen 
sulfide (H 2 S). In the case of guns where coppering of 
the bore takes place, the copper (and zinc, if gilding 
metal is used for rotating bands or bullet jackets) 
takes up essentially all of the sulfur, as described in 
Section 12.3.2. 

Pyrrhotite (FeS x ) has been found only rarely in 


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272 


THE CAUSES OF GUN EROSION 


eroded guns. 98 Experimental tests showed that iron 
sulfide was probably formed on steel test rods sub¬ 
jected to powder gases. 53 - 98 The results of these ex¬ 
periments to investigate the role of sulfur in gun ero¬ 
sion are described in Section 14.4. 

The results of the thermodynamic studies 60 led to 
the conclusion that FeS should be formed only in 
small quantities in portions of the gun tube not sub¬ 
jected to high temperatures since its formation is 
self-inhibiting due to the energy of reaction. This is 
true in spite of the fact that FeS is stable over a large 
range of conditions found in gun firing. 

Reactions Involving Nitrogen Compounds 

Probably the most puzzling question with respect 
to the chemical factors in gun erosion is that of the 
formation of the iron nitrides identified in eroded 
guns, which are described in Chapter 12. This subject 
has been discussed in some detail in Sections 14.5 and 
13.3.2. Extrapolation of existing data showed that 
iron nitrides could not be formed by reaction of the 
powder gases with the steel at the peak temperatures 
reached in firing of guns. 60 Experiments showed that 
iron nitrides are decomposed during firing. 28 From 
this it is presumed that the nitrides are formed in 
guns during the cooling stage. It is probably true, 
however, that nitrogen penetrates the bore surface at 
the peak pressures and temperatures especially if the 
surface is partially liquefied. The nitrides may form 
on cooling by exsolution from a nitrogen-rich phase 
or by reaction with the cooling gases. 

When the powder gases cool slowly, the concentra¬ 
tion of ammonia (NH 3 ) increases. Nitrides may form 
by reaction of the steel with this component of the 
gas. 28 The possibility of another reaction has also 
been proposed 138 —that of the steel with cyanogen 
(CN) or hydrocyanic acid (HCN). It is believed that 
the presence of these compounds containing both car¬ 
bon and nitrogen in the products of combustion of 
single-base powders may have considerable impor¬ 
tance, for complex iron cyanides have been found in 
cases where an oxide would be expected. 31 - 138 These 
compounds, which decompose at relatively low tem¬ 
peratures, have been detected in eroded guns as well 
as on test rods, as was mentioned in Section 12.2.2. 

When the concentration of ammonia in the gases 
increases, the concentration of hydrocyanic acid 
should also increase according to data for reaction 
( 11 ). 

2CO + NH 3 ^HCN + C0 2 + H 2 (11) 


A concentration of HCN was predicted in the cooling 
gases from a single-base propellant which might be 
sufficient to suppress oxide formation and extend the 
Fe 3 C field. 138 In addition the formation of nitrides 
and complex iron cyanides might result. 

Either carbide or nitride may be formed by the 
reaction of iron with cyanogen. Experiments were 
carried out to determine what factors were important 
in determining which compound would be formed. 138 
Mixtures of cyanogen and nitrogen were passed over 
iron filings at different temperatures. Depending on 
the conditions of the experiment, nitride or carbide 
or both were formed. It was found that the reactions 
involved are complicated. Temperature, CN/N 2 bal¬ 
ance, relative rates of diffusion of carbon and of nitro¬ 
gen, and the relative stabilities of cementite (Fe 3 C) 
and of the nitrides are all critical factors in deter¬ 
mining the end product. 

Conclusions 

In this section a number of reactions were discussed 
in the light of their possible or probable importance 
in the chemical attack of the bore surfaces of guns by 
powder gases. It cannot be stated definitely what re¬ 
actions occur during firing, for the mechanism of the 
chemical erosion is not thoroughly understood owing 
to lack of data on the reactions of iron with the con¬ 
stituents of propellant gases at the temperatures and 
pressures obtained in firing. Nevertheless, a great 
deal of information has been obtained, as summarized 
in Chapter 12, which shows that chemical reactions 
have taken place. 

In the case of guns fired with double-base powders, 
oxidation, with the formation of wiistite (FeO), has 
been observed. In the case of those fired with single¬ 
base powders, penetration of carbon and nitrogen, 
with the formation in many cases of carbides and 
nitrides, has occurred. The ultimate products found 
by thermodynamic calculations agree closely with 
those found in eroded guns. 

134 MECHANICAL FACTORS 

13,41 Powder Gases 

Removal of Material from Bore 

The final stage in the process of the erosion of steel 
guns is the removal of material from the bore surface, 
whereby the bore diameter is increased. This is effect- 


CONFIDENTIAL 



MECHANICAL FACTORS _ 273 


ed largely by the stream of powder gases propelling 
the projectile. Thus the powder gases not only heat 
the metal and react with it as described in the two 
preceding sections, but also sweep the surface. The 
rate and direction of removal of material depends on 
many factors. 

Loose iron oxide films and such solid products as 
result from reactions within the gases are readily re¬ 
moved. The ease of removal of metal or metallic 
products depends on whether they are liquefied at the 
bore-surface temperatures. Chemically altered (Sec¬ 
tion 13.3.2) steel liquefies at a lower temperature 
than steel; completely liquefied films are more mobile 
than “mushy” surface layers which may contain 
blades of cementite, as shown in Figure 8 of Chapter 
12. The higher the velocity of gas, the more effective 
is the removal of material. A bag-fired gun mentioned 
in Section 10.5.2. affoids a striking example of this 
relationship. In the chamber, where the velocity of 
the gas was low, there was much evidence of lique¬ 
faction due to the fact that the liquefied material had 
remained in place as shown in Figure 16 of Chapter 
10; whereas in the bore of the gun the liquefied 
material had been removed by the gas flowing at a 
higher velocity. 112 When velocities are increased by 
increasing the powder charge, removal of material is 
even more effective because the heat input to the bore 
has been increased and liquefaction takes place more 
readily. 

In general, the liquid and solid products are moved 
forward by the gas stream because of the strong 
shearing stress on the walls of the tube in that direc¬ 
tion. This effect may be modified by localized turbu¬ 
lent flow. Examinations of eroded guns have shown 
that peripheral hot gases emerging from the cartridge 
cases appear to strike the bore area in a nearly radial 
direction and eddy backward, building up in some 
cases thick deposits of once-molten material at the 
location of the mouth of the cartridge case. The im¬ 
portance of the thermal factors in this area is described 
in Section 13.2.5. The tongues of flowed material in 
the bores of guns are not always parallel to the bore 
axis. In the most severely eroded section of a gun 
showing exaggerated muzzle erosion the tongues 
pointed somewhat fanwise forward as from a local 
jet. 130 

Scouring Action 

The mechanical action of the propellant gas on the 
bore surface is aggravated by the presence of solid 


particles in the gas stream. Some of these, such as un¬ 
burned powder grains and barium and potassium 
salts may be directly attributed to the powder itself. 
In addition, the gases pick up from the eroded surface 
the hard, brittle products that have been formed by 
the reaction of the gun steel with the powder gases. 
These solid particles traveling with high velocity in 
the gas stream presumably exert an abrasive action 
on the surface of other parts of the gun bore. 

Gas Leakage 

Observation of Gas Leakage. Gas leakage is of prime 
importance as a cause of erosion only in special cir¬ 
cumstances, although the escaping of a relatively 
small amount of gas past the rotating band of the 
projectile because of imperfect obturation has been 
observed to be common. In connection with the bal¬ 
listic measurements with the 3-in. gun described in 
Section 4.2.2 an attempt was made to determine the 
time of ejection of the projectile by the cutoff of a 
beam of light passing across the muzzle and focused 
on a photo cell. The photocell circuit was interrupted 
prematurely by the opaque gas which escaped ahead 
of the projectile, as described in Section 4.3.20. Dur¬ 
ing the development of a gauge for measuring the 
acceleration of a projectile (Section 4.6), it was dis¬ 
covered that gas leakage occurred soon after the start 
of travel. High-speed photographs, made at Naval 
Proving Ground a number of years ago, of a large 
projectile emerging from the muzzle of a gun showed 
that a small amount of gas escapes ahead of the pro¬ 
jectile. 16 A method was developed for photographing 
the gas escaping ahead of a caliber .30 bullet when it 
was fired into an evacuated chamber. 49 A series of. 
pictures was made of the bullet at various positions 
in the bore. A cloud of gas in advance of the bullet 
was observed, but it was not possible to tell whether 
the gas leakage had occurred at the start of travel or 
later. 

Gas Leakage as a Cause of Erosion. A survey of the 
literature 16 showed that some early writers considered 
gas leakage one of the dominant causes of gun erosion, 
but the same survey also revealed that this theory 
was strongly contested. Most of the arguments, which 
were often based on observations of eroded vent 
plugs, were only speculative. The flow of the powder 
gas in a vent plug, however, is radically different from 
that in guns. Unless it can be shown that, when gas 
escapes past the projectile, it flows as through a vent 
the observations made on vent plugs should not be 


CONFIDENTIAL 




274 


THE CAUSES OF GUN EROSION 


applied to the formulation of a theory of gun erosion 
involving gas leakage. 

At present it is believed that the above conditions 
are met in certain cases in which localized scoring 
of the bore surfaces of badly eroded guns takes 
place. 479,541 Where such leakage does occur, relatively 
large quantities of molten metal are washed from the 
bore surface. The melting point of steel is easily 
reached because of the friction developed in the small 
passageway, as discussed in Section 5.4.4. The molten 
steel is rapidly washed away because of the very high 
velocity of the escaping gas. Thus gas leakage under 
certain conditions is an important cause of erosion 
from both a thermal and a mechanical standpoint. 

A theoretical treatment of the problem was under¬ 
taken which showed what critical conditions are ex¬ 
ceeded when gas escapes past a projectile as through 
a vent. 48 The danger of gas washing, according to this 
study, is greatest at the start of travel before the 
rotating band of the projectile has been engraved. 
One of the advantages of a pre-engraved projectile, 
therefore, is that a much smaller gas pressure is re¬ 
quired to start its movement, hence there is not as 
much opportunity for gas leakage to be effective. A 
description of some experiments with pre-engraved 
projectiles having different degrees of gas leakage is 
given in Section 31.4.2. 

Observation of Scoring. In some small- and medium- 
caliber guns that were examined with a microscope 
local scoring of the bore surface was observed near 
the origin of rifling. 112 The gouges produced by gas 
leakage or “blow-by” are illustrated in Figure 20 of 
Chapter 10. In some cases the channel evidently de¬ 
veloped by enlargement of certain cracks by the main 
stream of gas. Finally, it reached an optimum width 
to act like an erosion vent. 

In the case of large-caliber guns a channel for escap¬ 
ing gases may be present at the top of the bore before 
the projectile is fired because of the effect of gravity 
on its position in the bore. This may lead to severe 
scoring of the bore surface near the 12 o’clock position 
in the vicinity of the origin of the bore. Erosion of 
this type is especially noted in 155-mm guns. 216,284 

Gas leakage occurs and may be a serious cause of 
erosion at the forward joint of breech liners. The bore 
ahead of the liner is affected by gas washing. This 
type of failure occurs regularly in caliber .50 barrels 
containing stellite liners and is one of the factors that 
determines the life of such barrels. The practice of 
chromium plating the bore ahead of a stellite liner 
mitigates this effect but does not prevent it entirely. 


Gas Pressure 

Cracking of Bore. The pressure of the powder gases 
exerts a force on the walls of the gun tube. This is 
discussed in connection with the theory of tube stresses 
in Chapter 26. The stress due to gas pressure does not 
contribute directly to removing material but it does 
result in cracking of the gun bore which greatly facil¬ 
itates erosion, as is described in Section 13.5.3. 

Bore Expansion. Under conditions of rapid fire, the 
temperature of a barrel is built up with the result 
that the steel is weakened, and the barrel may be 
expanded by the gas pressure, with the resultant bad 
effects on ballistics described in Section 5.6.4. In 
the case of barrels containing erosion-resistant liners, 
expansion in the region of the forward end of the liner 
imposes a limitation on the length of burst. The meth¬ 
ods of mitigating this effect are described in Section 
24.2. 

Effect of Stress on Powder Gas Erosion. An experi¬ 
ment was carried out to determine whether the 
stressed condition of a metal surface was related to 
its erosion by powder gases, 103 using the apparatus 
described in Section 11.2.5. The measurements did 
not reveal any appreciable difference between the 
erosion of surfaces under tension and that of surfaces 
under compression, for the conditions of this experi¬ 
ment. However, even though the ballistics were regu¬ 
lated so that only a small amount of erosion would 
take place, plastic flow or melting appeared to be the 
predominant factor in the erosion, and thus the effect 
of stress on erodibility was rendered negligible by the 
larger effect of surface melting. 

Occlusion of Gases 

Hydrogen, which is an abundant constituent of 
powder gases, rapidly diffuses through steel at high 
pressures. It is conceivable that hydrogen diffusing 
through hot steel may react with the carbon in the 
steel to form methane, which would exert consider¬ 
able pressure within the steel. Thermodynamic cal¬ 
culations were made in order to determine if methane 
pressures could be a contributing factor in gun ero¬ 
sion. 49 From a general relation between the pressures 
of the gases and the free energy change of the re¬ 
actions, it was found that, if the gas in contact with 
the steel under a total pressure of 1,000 atm contained 
10% of hydrogen, bubbles of methane at a pressure 
in excess of 10,000 atm could form within the steel if 
the temperature were not more than 610 C. There is 


CONFIDENTIAL 



MECHANICAL FACTORS 


275 


no direct evidence, however, to suggest that such a 
process contributes to erosion. 

13 4 2 Projectile 

Band Pressure and Engraving Stresses 

The rotating band of the projectile exerts pressure 
on the gun tube as it moves through the bore. The 
stress exerted on the tube and the resulting deforma¬ 
tion is greatest in the vicinity of the origin of rifling 
because of the force necessary to engrave the band. 422 
The band pressure decreases toward the muzzle due 
to both wear and fusion of the band. This fact makes 
the design of the rotating band an important factor 
in the erosion of the gun that fires it. 273 When projec¬ 
tiles become tipped in the bore, so that engraving of 
the body takes place as described in Section 10.4.10, 
the erosion of the bore is asymmetric and extends to 
the muzzle, where it may be greater than elsewhere in 
the bore. 

The study of the stress-strain relationships during 
the process of engraving, which is discussed in Section 
7.3.5, led to the conclusion that the dominant stress 
in engraving is radial compression. Shearing in a ra¬ 
dial direction takes place at the edges of the lands. 
The greatest deformation usually occurs along the 
driving edges of the lands because of the force neces¬ 
sary to impart spin to the projectile. Once the rotat¬ 
ing band has been engraved, the band pressure is 
dependent on the size of powder charge, since band 
wear has been shown to decrease with the velocity of 
the projectile (Section 7.4.1). 

Abrasion 

The radial pressure between projectile band and 
bore produces friction which results in adding heat to 
the bore, as described in Section 5.2.2, and in abra¬ 
sion. The abrasion, it should be remembered, is not 
that of sorbitic steel but of the thermally and chem¬ 
ically altered steel at the bore surface. This material 
is not resistant to mechanical shock because, even 
though it is hard, it is brittle. After becoming heated 
in rapid-fire guns, however, the bore surface is soft 
and ductile; and swaging, as described in the next 
subsection, assumes more importance than abrasion. 
Loose material, such as oxide films, presumably are 
easily abraded from the bore surface. 

The abrasive effect can be minimized by decreasing 
the radial load, as discussed in Sections 27.3 and 27.4. 


In particular, when pre-engraved projectiles are em¬ 
ployed, erosion is markedly decreased, as is brought 
out in Section 31.4.3. 

Reduction of the coefficient of friction between the 
projectile and the bore by lubrication has little effect 
on the abrasion as long as conventional engraving- 
type projectiles are used. When the radial load is 
reduced, however, further improvement can be made 
by reducing the coefficient of friction between the 
projectile and the bore surface. Thus Parco-Lubrizing 
pre-engraved caliber .50 projectiles decreased the ero¬ 
sion of the lands, as shown in Figure 11 of Chapter 31. 

Swaging 

Careful microscope, horoscope, and star gauge ex¬ 
aminations of chromium-plated liners which were 
tested by firing in caliber .50 machine-gun barrels 
showed how the cracked chromium was removed by 
the swaging action of the bullets on the steel beneath 



Figure 9. Isothermal transformation diagram (S 
curve) for SAE 4140 steel. Curve C represents the start 
of transformation, curve D the end of transformation of 
austenite when cooled from the austenitizing tempera¬ 
ture and held at lower constant temperatures. The tem¬ 
perature represented by line A is the critical temperature 
for the alpha-gamma transformation (line GS in Figure 
5) of this particular steel and that represented by line B 
is the eutectoid temperature (line PSK in Figure 5) for 
this steel. Austenite is not stable below B and ferrite is 
not stable above A . (This figure is based on Figure 6 in 
NDRC Report No. A-300, which had been made avail¬ 
able by courtesy of U. S. Steel Corporation Research 
Laboratory.) 


CONFIDENTIAL 







276 


THE CAUSES OF GUN EROSION 


the chromium. 50 The steel of the lands during firing 
moved in three directions, as shown in Figure 21 of 
Chapter 10 and described in Section 10.5.3. 

In Section 13.2 the thermal transformation of the 
steel at the bore surface during firing was discussed. 
This change is very important in connection with the 
question of swaging. Temperature measurements 
made on barrels during long bursts of rapid fire to¬ 
gether with iS-curve data (Figure 9) indicate that the 
austenite had too little time to transform between 



(*-LAND 


Figure 10. Swaged land at 346 in. from origin of bore 
in caliber .50 aircraft barrel No. 1194, fired four 100- 
round bursts with complete cooling between bursts. 
25X. (This figure has appeared as Figure 8C in NDRC 
Report No. A-440.) 


rapidly fired rounds during the last part of a burst. 
The steel was in the soft austenitic condition when 
subjected to the impact of the projectile and was 
easily swaged. 59 Figure 10 shows a land of a machine 
gun barrel swaged down the middle as a trough. 

In the early rounds of a burst a hardened surface 
layer begins to form instantaneously during each 
cooling cycle at about 270 to 300 C, the temperature 
at which curve C in Figure 9 intersects the left-hand 
edge of the isothermal transformation diagram for 


SAE 4140 steel, which has a composition approxi¬ 
mating that of most gun steels. During the later 
rounds of a burst, when the minimum temperature 
for any one round exceeds 300 C, there is not time for 
the hardened layer to form before the next round is 
fired, because of the slowness of the transformation. 
As can be seen from Figure 9, transformation does 
not even begin above this temperature until several 
seconds have elapsed and then proceeds slowly (re¬ 
gion between curves C and D). 

Swaging is believed to be a general phenomenon 
but its effects are not always evident in nonplated 
guns because of the removal of metal by powder gas 
erosion. 

13 4 3 Design of Gun Bore 

The design of the gun bore as well as of the projec¬ 
tile is important in determining the stress that is 
applied to the tube. Rifling (Section 26.6) is essential 
in order to impart spin to the projectile but its pres¬ 
ence means that guns are structurally weak from the 
very start of their life because of the localization of 
stresses. Various modifications in the design of the 
bore have been carried out in order to mitigate ero¬ 
sion and increase accuracy and velocity-life as de¬ 
scribed in Sections 23.2.1 and 26.6.3. 

13 5 CRACKING OF THE BORE SURFACE 
13,51 Introduction 

Generally the most obvious surface feature of ero¬ 
ded bores is the cracked appearance. Cracks are not 
solely a cause of erosion nor are they solely an effect 
of erosion. Their relative importance with respect to 
cause and effect is not known. It can be correctly 
stated, however, that they always accompany erosion. 
Because of the complex nature of its role, cracking is 
treated here in a separate section. It is described first 
as a factor in aiding erosion. The formation of the 
cracks is then discussed. The development of certain 
types of crack patterns is described and illustrated in 
Section 10.5.1. 


13,5,2 Aid to Erosion 

Powder Gas Erosion 

Even if it had no other effect, the cracking of the 
bore surface would increase the rate of powder gas 


CONFIDENTIAL 




CRACKING OF THE BORE SURFACE 


277 


erosion simply because of the increased surface for 
reaction. The longitudinal cracks act as channel ways 
for the streaming gases and may become enlarged 
to the extent that localized gas leakage and resultant 
scoring may develop, as described in Section 13.4.1. 

Chromium-Plated Bore Surface. Cracking of the 
chromium plate is the underlying cause of failure in 
plated guns, 50 86 as discussed at length in Section 
20.2.1. Chromium under gun firing conditions is inert 
to chemical attack. A continuous coating of this 
metal protects steel barrels from chemical reaction 
with the powder gases although it does not, unless 
thick enough, as mentioned in Section 13.2.3, prevent 
thermal alteration of the steel. When the chromium 
plate is cracked due to firing stresses, however, the 
powder gases have access to the underlying steel and 
chemical reactions take place, especially if the steel 
is in the austenitic condition at the time. 124 The 
steel is eroded from beneath the chromium plate, and 
the latter, so undermined, is easily removed. 

Stress Erosion 

Pitting of the bore surface occasionally results 
when cracks intersect and small blocks of metal are 
torn away. A serious type of bore damage which for¬ 
tunately is not common to all types of guns is shear¬ 
ing of the lands. In some badly strained guns, cracks 
which start at the groove fillets may curve under the 
lands from both sides and eventually meet, where¬ 
upon whole sections of lands may be removed. 277 

13-5,3 Causes of Cracking 8 

Thermal and Transformation Stresses 

During the firing of a gun, the differentially con¬ 
fined bore surface is subjected to volume changes due 
to rapid heating and cooling and to the thermal 
transformation of steel (Section 13.2.3). The resulting 
strains set up in the surface layer may be relieved by 
cracking. The term “heat checking” has been em¬ 
ployed 16 to describe the fine crack system seen in the 
portions of eroded gun bores where no more than in¬ 
cipient melting of the surface has taken place, since 
thermally induced stresses would be expected to cause 
this type of shallow cracking. 

Not only the volume changes mentioned above but 
also the change in mechanical properties resulting 

g The study of the problem of cracking in gun tubes as man¬ 
ufactured during World War II was undertaken by the War 
Metallurgy Committee (Division 18, NDRC). 170 


from alteration of the bore surface affect the stress- 
strain relationships in the surface layer. The volume 
increase accompanying the inyersion to austenite was 
calculated from x-ray data on' eroded gun surfaces to 
be three per cent. 28 The resulting compressive stresses 
probably cause plastic deformation rather than crack¬ 
ing. Quench-cracking, however, may result when the 
bulk of the austenite is transformed to a brittle mar¬ 
tensitic layer during the rapid cooling by the mass of 
the gun steel beneath it. 124 In the case of single-shot 
guns, this type of layer is present before each succes¬ 
sive round. It is harder but less ductile than either 
austenite or the originally sorbitic steel and conse¬ 
quently is more readily fractured by the impact of the 
projectile. 

The thin, complex, outer white layer, discussed at 
length in Chapter 12, is partially liquefied during 
firing. It consists of materials only slightly subject to 
transformation stresses. On cooling it is brittle and 
tension cracks form in it which are closely in line with 
the cracks in the steel which the liquefied material 
has covered. 112 

The evidence which follows seems to show that, 
even though the thermal and transformation stresses 
may be of prime importance, heat alone is insufficient 
to cause the type of cracking referred to as “heat 
checking.” 

Studies of Thermal Shock by Electron 
Bombardment 

A study of the effects of thermal shock was made 
by subjecting gun steel specimens to electron bom¬ 
bardment in the presence of the inert gas argon, ac¬ 
cording to the procedure described in Section 11.3.2. 104 
Additional experimentation was carried out with ni¬ 
trogen or carbon monoxide substituted for argon. 
Metallographic examination showed that thermal al¬ 
teration had taken place; furthermore, analysis by 
electron diffraction showed that austenite was pres¬ 
ent in addition to ferrite. 137 Cracking had occurred 
only on specimens that had been bombarded in nitro¬ 
gen, but in no case was the crack pattern identical 
with that observed on eroded gun bores. This evidence 
shows that “heat checking” is not a characteristic 
effect of thermal shock alone. 

Causes of Cracking of Test Specimens 

A testing method was devised in order to determine 
the resistance of metals to cracking. 51 The polished 


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278 


THE CAUSES OF GUN EROSION 


surface of a block of the metal to be tested was ex¬ 
posed to a stream of powder gases in a small explosion 
vessel, described in Section 11.2.4. The surface was 
then examined under a microscope. 

Cracks that could be attributed to mechanical 
stresses were observed. In addition, a crack system 
similar to the so-called “heat checking’' in eroded 
guns was found to start on steel specimens at the 
place where the hot gases had first come in contact 
with it and with successive firings to progress “down¬ 
stream.” This type of cracking was not found on an 
exposed molybdenum surface. 



Figure 11. Eroded surface of 3-in. gun liner No. 1460 
at the location of the mouth of the cartridge case. The 
pattern on the eroded area was influenced by tool marks. 
20X. (This figure has appeared a? Figure 8A in NDRC 
Report No. A-440.) 

The fine crack system did not appear immediately 
in the case of steel blocks; yet it was apparently re¬ 
lated to thermal transformation, which, however, 
could not be the sole cause of cracking, since the 
latter did not start until after several firings. This 
may mean that time was required for sufficient 
chemical alteration to aid in the formation of the 
cracks. 

Further experiments in which black powder or its 
components were added to the charge strengthened 
this idea. The hypothesis was suggested, therefore, 
that “heat checking” is caused by a combination of 
thermal and chemical factors. This bears out the con¬ 


clusion from the electron bombardment experiments 
just described that “heat checking” is not produced 
solely by thermal shock. 

Causes of “Pebbling” in Guns 

Pebbling and heat-checking are terms used to de¬ 
scribe the characteristic appearance of an eroded gun 
bore in which general melting has not occurred; the 
latter is descriptive of the pattern, whereas the for¬ 
mer denotes also the doming of the units of the pat¬ 
tern by liquefaction and erosion as described in Sec¬ 
tion 10.5.2and shown in Figure 16 of Chapter 10. 112,124 



Figure 12. Elongated pits resulting from fluxing of 
inclusions on groove surface of 3-in. gun liner No. 1460 
at 1 in. from origin of bore. 25X. (This figure has ap¬ 
peared as Figure 6C in NDRC Report No. A-440.) 

Since the cause of this type of cracking has not been 
found, the use of the term pebbling is preferable in 
that it does not imply a cause. 

The hypothesis mentioned above, namely, that this 
phenomenon is caused by a combination of thermal 
and chemical factors, is in agreement with the results 
of the metallographic examination 124 of the liners 
used in the powder testing program described in 
Chapter 15. No pebbling was observed with a small 
number of rounds with a single-base powder although 
thermal alteration took place in the first round. The 
conclusion drawn from this fact was that chemical 
alteration was necessary before melting of the surface 
could take place. Furthermore, it is believed that 
enough carbon or nitrogen or both has to penetrate 
the steel before austenite formed during firing can be 


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CRACKING OF THE BORE SURFACE 


279 





Figure 13. Extent of progressive stress-damage in 76-mm M1E5 gun tube No. 2128 after 2,000 rounds. Arrow points 
to driving side of land. Specimens were nickel-plated prior to polishing. (A) y 2 in. ahead of origin of rifling; (B) 9 in. 
ahead of origin. (This figure has appeared as Figure 7B and 7C of Watertown Arsenal Laboratory Report No. WAL 731 /95.) 


from which cracks may develop in further firing. 112 

Progressive Stress-Damage 

Progressive stress-damage, a term coined at Water- 
town Arsenal, is defined as “all phenomena associated 
with the initiation and propagation of cracks in guns 
as a result of the mechanical stresses imposed by re¬ 
peated firing.” 277 The mechanical stresses are those 
due to engraving of the rotating bands or of the pro¬ 
jectiles themselves, band pressure, and powder gas 
pressure, all of which are discussed in Section 13.4. 
Progressive stress-damage cracks are found at places 
where there were stress raisers. The most serious pro¬ 
gressive stress-damage occurs as deep cracks at the 
groove fillets. An illustration of the extent of progres¬ 
sive stress-damage is shown as Figure 13. The prop¬ 
agation of these cracks may eventually cause failure 
of the tube by rupturing or by shearing of the lands. 
The extent of progressive stress-damage in a number 
of different worn gun tubes, 274 ’ 275 - 277 ’ 278 - 280 - 281 especi¬ 
ally those for the 76-mm gun, Ml, 264 - 265 - 266 - 267 - 269 has 
been studied in recent years at Watertown Arsenal. 


stabilized. Quenching cracks are inevitable if more 
carbon than the eutectoid composition is dissolved in 
the austenite. It would seem then that pebbling may 
be the result of a combination of quench-cracking 
and incipient metlting of the surface which tends to 
draw itself into globules. 

Stress Raisers 

Tool marks have been considered a cause of the 
cracking in gun tubes when firing stresses are applied, 
but objections to their importance have been raised 
in that there are many more cracks in eroded guns 
than there were tool marks to start with. 16 Tool 
marks are stress raisers, however, and as such can 
influence the initiation and direction of cracks as 
shown in Figure ll. 112 (This figure is also a good il¬ 
lustration of pebbling.) 

Inclusions in the steel at the bore surface also in¬ 
fluence cracking. This may be a cause of failure in 
guns made of a dirty steel. 262 The inclusions react 
with the powder gases and may “explode out” leav¬ 
ing an elongated pit such as those shown in Figure 12 


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PART IV 

EROSIVE ACTION OF PROPELLANTS 



I were better to be eaten to death with a rust 
than scoured to nothing with perpetual motion. 

—William Shakespeare 
“King Henry IV” 


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Chapter 14 

EFFECTS OF CONSTITUENTS OF THE POWDER GASES ON GUN STEEL 

By W . D. Urry a 


141 INTRODUCTION 

T he investigations described in this chapter 
were largely complementary to the general study 
of the nature of eroded bore surfaces, described in 
Part III, but were intended at the same time to serve 
a further purpose. It was hoped that they would elu¬ 
cidate the particular role that each of the components 
of a propellant gas, treated individually insofar as 
feasible, might play in erosion or in the production of 
conditions favoring erosion by a more general mech¬ 
anism. It was even considered at the outset quite 
possible that such studies would indicate the manner 
in which the composition of propellants then in use 
might be modified, perhaps only in a minor respect, 
in order to mitigate erosion. 

The investigations have not demonstrated that the 
erosion of steel guns can be startlingly diminished by 
a practical modification of the present powders. In¬ 
stead they have contributed to the evidence already 
presented in Chapter 12 that carburization, oxida¬ 
tion, nitriding, reactions of sulfur gases, or possibly 
the direct removal of iron as a volatile compound like 
iron carbonyl, present almost insurmountable diffi¬ 
culties if a steel bore surface is to be considered, es¬ 
pecially under the conditions of the upswing in ballis¬ 
tic level demanded by modern artillery practice. 

The above reactions have been carefully studied 
under ordnance conditions or in experiments under 
laboratory conditions simulating one or more of the 
conditions pertaining to a gun. The experiments may 
be divided into two general groups in another sense. 
In the one group the individual constituents of a 
powder gas were used; in the other the powder gases 
were employed collectively while studying the role of 
one particular constituent or element. Only in the 
latter case, of course, can there be true ordnance con¬ 
ditions. 

The constituents of the powder gases that have 
been studied in relation to reactions with gun steel 
include carbon monoxide and carbon dioxide, nitro¬ 
gen, hydrogen, and hydrogen sulfide, some of these 


a Physical Chemist, Geophysical Laboratory, Carnegie In¬ 
stitution of Washington. 


gases having been investigated individually, others 
in various combinations. In other experiments em¬ 
phasis was placed on a particular element, such as 
carbon, nitrogen or sulfur, while the gases were pres¬ 
ent collectively, being obtained from a standard pro¬ 
pellant. In this case the minor constituents such as 
methane and ammonia were present as well as water 
vapor, which was not otherwise investigated. Black 
powder, used with the primer, may constitute no 
negligible fraction of a gun charge. Except in small 
arms, the sulfur in the black powder is the chief 
source of the sulfur gases and therefore a study of the 
reactions with the sulfur gases introduced the ques¬ 
tion of the effects of the other constituents of black 
powder, namely, potassium nitrate and charcoal, to 
which some attention was given. 

The chemical factors in the causes of gun erosion 
and their interrelation with other factors are sum¬ 
marized in Chapter 13. A great many of the conclu¬ 
sions concerning them were derived from or strength¬ 
ened by the results of the above experiments. A gen¬ 
eral review of the results of these investigations in¬ 
dicates that many of the undesirable features of the 
reaction of the powder gases with gun steel can be 
attributed to the presence of the carbon gases, con¬ 
cerning which little can be done beyond an attempt 
to adjust their proportions or otherwise attain certain 
conditions of balance between carburizing and oxidiz¬ 
ing states. Nitrogen, perhaps present as the activated 
molecule or as ammonia, plays its role by forming- 
nitrides. It is difficult to imagine its removal from the 
propellant. Investigation of the part played by hy¬ 
drogen is very incomplete. The effects of water vapor, 
as an individual constituent of the powder gas, have 
never been studied. There is indirect evidence that 
its role may be subsidiary because the erosion of vent 
plugs by artificial gas mixtures free from water vapor 
is the same order of magnitude as that caused by 
powder gases derived from standard propellants. 
Sulfur definitely increases erosion, probably in all 
cases except where direct melting of the steel is occur¬ 
ring. The exact mechanism whereby the presence of 
a few tenths of one per cent of hydrogen sulfide can 
contribute appreciably to the erosion is not com¬ 
pletely understood but experiments in three labora- 


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283 



284 


EFFECTS OF CONSTITUENTS OF POWDER GASES 


tories show that the elimination of sulfur, largely 
present in the black powder, lessens the erosion. 

Many of the reactions studied, and also the reac¬ 
tion velocities, are strongly favored by increasing the 
temperature, and insofar as any of these reactions 
may bear some quantitative relation to the erosion 
that it causes, either directly or indirectly, the sub¬ 
stitution of cooler propellants is worthy of consider¬ 
ation. However, there are some reactions, like the 
formation of iron nitrides and the oxides Fe 3 0 4 and 
Fe 2 0 3 , and possibly the production of iron carbonyl, 
that are strongly favored by lowered temperature. 
Nevertheless, the substitution of new propellants of 
lower flame temperature but with unimpaired im¬ 
petus, for example, the cool albanite and RDX pow¬ 
ders, 11 should lead to a mitigation of erosion. Studies 
of the erosiveness of some of the RDX powders are 
described in Chapter 15. The slight modification of 
the older, standard propellant is discussed further in 
Section 15.6. 

14 2 REACTION OF CARBON GASES WITH 
GUN STEEL 

14,2,1 Introduction 

Penetration of carbon into the bore surface of guns 
in both the sense of the formation of carbon com¬ 
pounds and the penetration of the steel may be a 
factor in erosion. The melting point of the surface 
material may be lowered from 1450 C to 1135 C by 
the presence of 4.3% of carbon, as shown in Section 
13.3.2. In some experiments with a rifle barrel this 
amount of carbon was exceeded in nearly all cases 
close to the surface and near the origin of rifling." 

Previous carbon analyses 481,490 are inadequate and 
interpretation is complicated by the use of completely 
eroded barrels in which the added carbon may be 
solely contained in carbonaceous material filling the 
extensive fissure system, as described in Section 
12.4.1, and coating the bore surface. In order to pur¬ 
sue the problem further, it was deemed necessary to 
develop a technique wherewith the presence of pro¬ 
pellant carbon in layers measured in fractions of 
microns could be detected and measured. It was felt 
that attention should be confined mainly to the effect 
of a single round in an unproofed barrel. 

Calculations of the probable bore-surface temper¬ 
ature of a caliber .30 barrel indicated that the trans- 


b Described in Chapter 6 of Volume 1, Division 8. 


ition of the iron from the alpha to the gamma phase 
does not take place when fired single-shot with the 
barrel at ambient temperatures. Therefore any exten¬ 
sive penetration of carbon into the steel is unlikely. 
In guns of larger bore this transition does occur, as is 
brought out in Section 13.2.3. Calculations showed 
that temperatures, and to some extent, the significant 
times involved, pertaining to guns up to 90-mm bore, 
could be simulated in suitably preheated caliber .30 
barrels. 

14 2 2 Summary of Recent Results 

Experiments with preheated caliber .30 barrels 
showed that carbon from propellant gases can pene¬ 
trate the bore surface of a gun to a marked degree in 
two senses. At the immediate surface there are carbon 
compounds of iron and possibly other elements form¬ 
ed. Below these surface layers carbon penetrates into 
the steel. These characteristics were found even after 
firing a single round in unproofed preheated caliber 
.30 barrels. 

Carbon Compounds. The average content of pro¬ 
pellant carbon in the reaction products is essentially 
independent of the bore-surface temperature and of 
the number of rounds. The distribution of the pro¬ 
pellant carbon within this layer is affected by the 
bore-surface temperature, for example, the carbon 
content of the part of the layer nearest the surface 
decreases with increasing bore-surface temperature. 
A particular feature within the layer of reaction prod¬ 
ucts is the universal occurrence, in these experiments, 
of a subsurface layer containing more propellant car¬ 
bon than a layer nearer to the surface that is often 
almost devoid of carbon. This is in accord with the 
lamellar structure of the layer of reaction products 
described below and the bottom of this buried, car¬ 
bon-rich layer seems to define the boundary between 
reaction products and steel. 

The thickness of the layer of reaction products, for 
a single round, is almost independent of the bore- 
surface temperature, particularly near the origin of 
rifling. The thickness of this layer was increased ap¬ 
preciably by firing more rounds. 

A change of propellant causes an important change 
in the carbon content of the reaction products. A cool 
RDX powder yielded reaction products that con¬ 
tained much less carbon than the products from IMR 
powder at the same bore-surface temperature. The 
products with a hot, double-base powder contained 
very little carbon near the origin of rifling but a con- 


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REACTION OF CARBON GASES WITH GUN STEEL 


285 


siderable amount, although at a very shallow depth, 
towards the muzzle, by comparison with IMR pow¬ 
der. The average content of propellant carbon and 
original carbon in any of the tests could produce at 
the most 30% of cementite in the layer of reaction 
products, and the introduction of carbon is therefore 
not the sole cause of the chemical alteration of the 
bore surface; in fact, with double-base powder it ap¬ 
pears to be but a very minor cause. 

Penetration into the Steel. There is very little pen¬ 
etration of carbon into the steel below a peak bore- 
surface temperature of about 900 C. Above this tem¬ 
perature the depth of penetration increases rapidly 
and the total amount of carbon that has penetrated 
increases uniformly with increasing temperature. In 
the caliber .30 barrel there was no evidence of pen¬ 
etration into the steel beyond 20 calibers forward of 
the origin of rifling. 

A change of propellant apparently causes no change 
in the depth of penetration provided the CO/C0 2 
ratio in the powder gases is a carburizing one. If the 
CO/CO 2 ratio is an oxidizing one and the bore-surface 
temperatures are high, which is the case with FNH- 
M2 powder, there is no penetration of propellant 
carbon into the steel. The dependence of CO/C0 2 on 
temperature and on composition of powder is dis¬ 
cussed in Sections 2.3.4 and 13.3.3. 

Relation to Commercial Carburizing. It is very diffi¬ 
cult to draw any parallels between carburization 
phenomena in a gun barrel and observations concern¬ 
ing carburization of steels. The difficulties are man¬ 
ifold : firstly, a lack of understanding of the phenom¬ 
ena in the immediate surface layer of a steel under¬ 
going commercial carburization; secondly, the violent 
fluctuations in a gun of all the variables that control 
carburization. Nevertheless, the initial slopes of the 
carbon depth curves just below the interface between 
reaction products and steel give diffusion coefficients 
comparable with values reported in the literature. 
Carburizing temperatures corresponding to these dif¬ 
fusion coefficients are in reasonable agreement with 
the theoretical bore-surface temperatures. 

Nature of Reaction Products. The nature of the 
reaction products, concerning which the use of the 
method to determine carbon content offered no clue, 
was studied by electron and x-ray diffraction. 137 The 
cementite (Fe 3 C) formed by the firing of IMR pow¬ 
der was separated into a layer removable with nital 
(5% HNO 3 in alcohol) and a film that adhered to the 
steel. A very thin film on the surface of the reaction 
products contained magnetite (Fe 3 0 4 ). In a qualita¬ 


tive manner, this analysis and the order in which the 
compounds appear correspond well with the distribu¬ 
tion of carbon in the layer of reaction products. 

The cause of the separation of cementite into two 
parts, the one removable, the other adherent, appears 
to be the presence of a noncarbon compound separat¬ 
ing the two zones. Evidence for the existence of this 
separating film is found in the carbon penetration 
curves. In other work 31 on the nature of the reaction 
products on blocks of steel exposed to the gases from 
a single-base powder, the nature of the compounds 
and the order of their occurrence were the same ex¬ 
cept that in this case there was evidence for the for¬ 
mation of a thin film of a complex cyanide. The cover¬ 
ing of magnetite was particularly a feature of highly 
heated blocks; a decrease in the carbon content of the 
immediate surface with increasing initial temperature 
of the preheated barrels is indicative of the formation 
of increasing amounts of a noncarbon compound. 
Carbon penetration curves for the hottest barrels 
suggest some decarburization following carburization. 
It is possible that a complex cyanide is momentarily 
formed but that most of this is lost by oxidation to 
magnetite. The surface carbon content is high enough 
in the cold barrel to permit the presence of a cyanide 
but the nature of the products was not studied. 

14,2,3 Experimental Determination of 
Carbon Penetration" 

Method 

All the carbon components of a powder gas in the 
act of propelling a projectile under standard ordnance 
conditions were tagged with radioactive carbon. The 
barrel was cut up into short lengths after firing and 
the resulting tubes were mounted in turn to form 
the hollow-cylindrical cathode of a Geiger-Mueller 
counter, described in Section 11.5.5. This is an appa¬ 
ratus that measures the passage of any radioactive 
ray traversing the cylindrical space, in particular the 
beta-rays emitted from radiocarbon in the surface of 
the cathode. 

Thin layers of the bore surface were removed suc¬ 
cessively by electropolishing for short periods. 22 The 
thickness of the layer so removed, controllable from 
0.2 micron upwards, was determined by analyzing 
the electro polishing solution for iron. 21 By alternately 
electropolishing and measuring the activity of the 
remaining radiocarbon, the content of propellant car¬ 
bon in the successive layers was indexed by the fall in 
activity. 


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286 


EFFECTS OF CONSTITUENTS OF POWDER GASES 


With respect to radioactivity techniques a novel 
feature of this method is the very efficient use of the 
specimen, the bore surface, by making it a part of the 
counter itself. A novel feature as far as analysis is 
concerned is that preservation of the “cut” for carbon 
analysis is unnecessary. 

Investigation was confined to chromium-molybde¬ 
num WD 4150 steel as specified by the U. S. Army 
for caliber .30 barrels. A special mount for the caliber 
.30 assembly was necessary in order to fire from ini¬ 
tial temperatures up to 700 C. The barrels were pre¬ 
heated by a coil of Nichrome wire wound directly 
onto a barrel and the bore surface was protected dur¬ 
ing the preheat period by the passage of nitrogen 
especially purified to remove oxygen. (See Figure 18 
in Chapter 11.) 

Validity of Assumptions 

Experiments were performed that showed: 

1. That the dissociation of the barium carbonate 
carrying the radiocarbon in the charge was complete. 

2. That there was equipartition of this radiocar¬ 
bon between the carbon monoxide and dioxide so that 
it was immaterial whether carburization and forma¬ 
tion of reaction products proceeded from one or the 
other. 61 (This phase of the work has been discussed in 
Section 2.3.5.) 

3. That the normal “ballistic level” (muzzle ve¬ 
locity and maximum pressure) is unaffected by the 
addition of 10 to 30 mg of barium radio carbonate. 

4. That the various experiments were conducted 
at the same ballistic level for IMR powder and at as 
nearly the same as possible for FNH-M2 and RDX 
powders. 

Limitation on Interpretation of Results 

The results must be interpreted in the light of the 
following limitations: 

1. Measurements were made on surfaces which 
extended over a 2-in. (6.7 calibers) length of barrel. 
Consequently areal differences in carbon content 
were averaged out to this extent. Apart from longi¬ 
tudinal variations a very real difference between lands 
and grooves was indicated. 

2. Carbon concentrations and depths were calcu¬ 
lated on the basis of the weight fraction of iron in the 
steel. Consequently, in that portion of a carbon-depth 
curve representing reaction products, the carbon con¬ 
tent is probably higher than the true value, but not 


by more than a factor of 1.2. The depth is correspond¬ 
ingly too shallow. 

3. The radiocarbon measures only the presence of 
carbon acquired from the propellant gases. If there is 
no exchange, that is, simultaneous carburization and 
decarburization, the total carbon content is given by 
adding the original carbon (nominally 0.5%) to the 
amount measured by the tracer and presented in the 
graphs. 

4. Smooth curves are plotted but actually “cuts” 
of finite thickness were taken. While a sudden change 
in the carbon content of a very thin layer will show 
in the analysis, it may not be in true perspective. 

5. The abnormal distribution of initial tempera¬ 
ture extended that length of bore surface forward of 
the origin of rifling which was momentarily raised to 
about the same peak temperature. Carburization ap¬ 
peared to extend further towards the muzzle than it 
probably does in actual practice. 

14 2 4 Bore-Surface Temperatures" 

The importance of the bore-surface temperatures 
with particular reference to the temperature of the 
alpha-gamma transition as controlling carbon pene¬ 
tration has already been mentioned. 0 This tempera¬ 
ture is roughly 750 C although there is abundant 
evidence that a considerably higher temperature is 
necessary to form an altered layer in guns, probably 
because of the ultra-short times. 59 Carbon penetra¬ 
tion was also found to cease far short of the 750 C 
isotherm below the surface. 

The instantaneous bore-surface temperatures can 
at present be obtained only from theoretical consider¬ 
ations by the methods described in Section 5.4.1. A 
few results of a large number of such calculations, 
including temperatures at depth, having a bearing on 
the problem of carburization are presented here. Fig¬ 
ure 1 illustrates the extent to which a preheated cal¬ 
iber .30 barrel simulates the temperature-time rela¬ 
tions in a 3-in. gun. The maximum peak temperature 
reached in the caliber .30 barrel starting at 27 C is 
computed to be about 630 C and obviously the part 
played by carbon in the erosion of caliber .30 rifle 
barrels in normal usage must be negligible; in fact, 
there is no erosion problem. 

The bore-surface temperatures for four firing tests 
with a preheated caliber .30 barrel used to determine 

c The formation of reaction products on the surface is con¬ 
trolled by other temperature considerations not connected 
with this transition, as discussed in Section 13.3.3. 


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REACTION OF CARBON GASES WITH GUN STEEL 


287 



Figure 1 . A comparison of the calculated bore-surface 
temperatures near the origin of rifling for a 3-in. AA 
gun, M3 and a preheated caliber .30 rifle. Curve A— 
3-in. gun initially at 200 C, e.g., after rapid fire; Curve 
B—3-in. gun initially at ambient temperature; Curve C 
—caliber .30 rifle, M1903A1 preheated to 450 C. (This 
figure was based on Figure 9 of NDRC Report No. 
A-427.) 



Figure 2. A comparison of bore-surface temperatures 
near the origin of rifling for a 3-in. AA gun, M3 and pre¬ 
heated caliber .30 barrels used in various tests for carbon 
penetration. Curve A—the 21st round of a burst in a 
3-in. gun with NH-M1 powder at 2,750 fps and 20 
rounds per minute; Curve B—the first round of the 
burst in the 3-in. gun; Curves C—caliber .30 barrels at 
various initial temperatures. (This figure appeared as 
Figure 11 in NDRC Report No. A-427.) 


carbon penetration are compared in Figure 2 with 
those for a 3-in. gun at the beginning and end of a 
burst. 106 Figure 3 shows the increase in bore-surface 
temperature due to a change of propellant from IMR 
to FNH-M2 in one of the firing tests. Figure 4 shows 
that the carbon penetration with a cool RDX powder 
and the hotter IMR powder can be compared when 


the former is fired at a higher initial temperature to 
compensate for its lower flame temperature. On firing 
at the same initial temperature, the RDX powder 
would produce lower bore-surface temperatures and 
carburization would be lessened. 

The decrease in carburization to be expected in. a 
90-mm gun by substituting a cool RDX powder for 
NH-M1 propellant on the basis of these experiments 
was investigated. It was estimated that the total 
depth of penetration and the total amount of carbon 
entering the bore surface might be reduced by at 
least 50%. The layer of reaction products would con¬ 
tain less carbon, but its thickness would apparently 
be about the same as with NH-M1 propellant. 



Figure 3. A comparison of bore-surface temperatures 
near the origin of rifling in a preheated caliber .30 rifle 
for single-base (IMR) and double-base (FNH-M2) 
propellants. The muzzle velocity was 2,685 fps in each 
case. (This figure was based on Figure 12 of NDRC 
Report No. A-427.) 



Figure 4. The similarity in the temperature-time rela¬ 
tions of the bore surface near the origin of rifling in two 
firing tests, one with an IMR powder, the other with an 
RDX-CC1 powder fired at a higher initial temperature 
to compensate for its lower flame temperature. (This 
figure appeared as Figure 15 in NDRC Report No. 
A-427.) 


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288 


EFFECTS OF CONSTITUENTS OF POWDER GASES 


14 2 5 Carbon Penetration" 

Measurements of carbon penetration into a caliber 
.30 barrel initially at room temperature exhibited 
little more than a veneer of reaction products close to 
the origin of rifling, although an extremely thin film 
in the immediate surface had the very high carbon 



2 4 6 8 10 

DEPTH IN MICRONS 


Figure 5. The penetration of carbon from propellant 
gases derived from a single-base powder (IMR), into the 
bore wall of a preheated, unproofed caliber .30 steel rifle 
barrel. The inset (upper right) shows the initial tem¬ 
perature of the barrel at different places along its axis 
and the positions of the specimens. In this figure, and in 
Figure 6, the curves are actually continuous but the 
scale for carbon content has been enlarged at an arbi¬ 
trary depth. The dashed curve for section A shows the 
small carbon content of the reaction products and lack 
of any penetration with a double-base propellant (FNH- 
M2) in another barrel (see Figure 3). 


content of 19.3%, indicating possibly a complex cyan¬ 
ide. A typical example of the carbon penetration 
close to the origin of rifling from firing a single round 
of IMR powder is to be seen in Figure 5. Carbon pen¬ 
etration was always investigated in two further sec¬ 
tions extending to 11 in. forward of the origin of 
rifling but the penetration was always negligible, 
being less than 1/x deep and showing none of the 
peculiarities that occur near the origin of rifling. Thus 
extensive carbon penetration occurs only in the region 
of severest erosion. 

Effects of Temperature 

The total depth of carbon penetration increases 
with increasing temperature of the bore surface. The 
thickness of the layer of reaction products, however, 
remains nearly the same for a single round. The con¬ 
tent of propellant carbon in the immediate surface 
decreased from 19.3 to 3.7% in four firings in which 
the initial temperature was increased from 25 to 
565 C. 

Effect of Number of Rounds 

The firing of ten rounds using IMR powder with 
radiocarbon in each round caused a considerable 
thickening of the layer of reaction products, particu¬ 
larly at a point about 3 calibers forward of the origin 
of rifling; but the depth of penetration into the steel 
beneath was the same as for a single round. It appears 
that a given carbon-depth relation is re-established 
with each round, as might be expected if penetration 
is chiefly a function of the temperature-time relations 
during firing. While the depth of penetration into the 
steel was unchanged, the propellant carbon content 
was increased in the surface of the steel below the 
reaction products from 0.2 to 1.9%. 

Effect of Propellant 

The principal carbon penetration experiments were 
conducted with a single-base powder of the IMR 
type. The examination of bore surface materials de¬ 
scribed in Chapter 12, however, had indicated that in 
the case of Service guns the type of powder caused a 
difference in the chemical nature of the erosion prod¬ 
uct, cementite (Fe 3 C) being associated with the firing 
of single-base powder and wiistite (FeO) with double¬ 
base. 

Furthermore, thermodynamic calculations 60 (Sec- 


CONFIDENTIAL 
















REACTION OF CARBON MONOXIDE WITH GUN STEEL 


289 


tion 12.2.4) indicate that for the bore-surface tem¬ 
peratures shown in Figure 3 and the C0/C0 2 ratio 
in the gases from a double-base powder, FeO would 
be the ultimate reaction product. Similar calculations 
for IMR powder give cementite as the equilibrium 
product. On the expectation that there might be a 
difference in the degree and extent of the carbon pen¬ 
etration between firings with single- and double-base 
propellants, experiments were conducted with both 
FNH-M2 and RDX powders. 

FNH-M2 Powder. A portion of the results obtained 
in the experiment with FNH-M2 powder are given in 
Figure 5 for the region near the origin of rifling. A 
search for carbon to a depth of 8 \x revealed none 
beyond l/z. Thus carbon plays an insignificant role 
in the reaction products remaining on the surface 
with FNH-M2 powder and fails to penetrate the steel 
at all. 

RDX Powder. This powder had the same nominal 
composition as that used in one of the erosion tests 
at the Franklin Institute. It caused negligible erosion 
under certain conditions described in Table 3 of 
Chapter 15. The bore-surface temperature and the 
CO/CO 2 ratio are such that chemical thermodynam¬ 


ics would predict a carburizing action. In Figure 6, 
where a comparison of the carbon penetration is 
made between IMR and RD^C powders for the same 
bore-surface temperature (see Figure 4), it can be 
seen that the depths of penetration are the same, 
although the average content of propellant carbon in 
the layer of reaction products with RDX powder is 
only about one-half of that with IMR powder. Other 
reaction products not containing carbon are presum¬ 
ably present to a greater extent with the cool RDX 
powder than with IMR powder. This is in accord 
with the identification of iron nitride (listed in Table 
12 of Chapter 15) on the bore surface of a steel barrel 
that had been fired with this powder in the caliber 
.50 erosion-testing gun, 123 as described in Section 
15.3.4. 

14 3 REACTION OF CARBON MONOXIDE 
WITH GUN STEEL 

14,31 Introduction 

The possibility that the formation of carbonyls of 
iron might play a role in gun erosion has been care- 



Figure 6. The penetration of carbon from propellant gases derived from RDX-CC1 powder into the bore wall of a 
preheated, unproofed caliber .30 steel rifle barrel. These curves are compared with those for the hotter IMR powder at 
lower initial temperature but same bore-surface temperatures (see Figure 4). The inset (upper right) shows the initial 
temperatures of the barrels and the positions of the specimens. (This figure appeared as Figure 30 in NDRC Report 
No. A-427.) 


CONFIDENTIAL 















290 


EFFECTS OF CONSTITUENTS OF POWDER GASES 


fully studied. 19 - 62 - 63 So far only laboratory experi¬ 
ments have been performed. Although a positive 
identification of a volatile carbonyl under exact ord¬ 
nance conditions has not been obtained, the work is 
being continued 11 with this ultimate end.in mind. The 
preliminary studies that can be reported here fall into 
two groups: measurement of the weight losses of ero¬ 
sion vent plugs subjected to gases resulting from ex¬ 
plosions of carbon monoxide-oxygen mixtures and 
the recovery and identification of volatile iron com¬ 
pounds. 

Explosion of the above mixtures produced a mix¬ 
ture of carbon monoxide and dioxide with a small 
amount of free oxygen due to the dissociation of car¬ 
bon dioxide. In addition to these gases there were 
present in the eroding gas mixture the impurities in 
the original carbon monoxide in either a free or com¬ 
bined state. These consisted of 4 to 5% of nitrogen 
and 2 to 3% of hydrogen. Especial care was taken to 
remove the iron carbonyl prevalent in commercial 
tanks of carbon monoxide. An increase in the oxygen 
content increased the flame temperature and changed 
the composition of the eroding gases but the explo¬ 
sion could be modified by the simultaneous addition 
of carbon dioxide to retain the same composition 
while lowering the flame temperature by any desired 
amount. 

14 3 2 Summary of Results 

The erosion of vent plugs made of gun steel with a 
34-in. bore diameter was found to -be strongly depen¬ 
dent on the flame temperature of eroding gases pro¬ 
duced by carbon monoxide-oxygen explosions. The 
erosion showed little if any dependence on the 
C0/C0 2 ratio in the eroding gases. The magnitude of 
the weight losses and the lack of any dependence on 
this ratio indicates that the steel per se was melted 
under these experimental conditions, except possibly 
at the lowest flame temperature. The C0/C0 2 ratios 
and the character of the reaction products seem to 
correspond to the conditions for a double-base pro¬ 
pellant containing 20% or more of nitroglycerin. 

The addition of small amounts of hydrogen sulfide 
substantially increased the erosion, particularly with 
the cooler bore surfaces; but when the erosion was 
extremely severe this effect was relatively small. Hy¬ 
drogen sulfide is a catalyst for the synthesis of iron 


d At the Johns Hopkins University under a contract with 
the Army Ordnance Department. 


carbonyl and increase in erosion might have been 
caused by the greater production of this volatile 
compound; but there are other equally feasible ex¬ 
planations. 

Iron carbonyl could not be detected in the gases 
collected after passage through vent plugs. When the 
bore-surface area was increased by substituting cal¬ 
iber .30 barrels for the vent plugs, iron carbonyl was 
detected to the extent of a few tenths of 1 per cent of 
the total iron deposits formed during the explosion 
of the coolest gas mixture or a solid propellant. The 
major portion of the iron occurred in iron-bearing 
deposits in a muzzle tube designed to collect the gases 
and as a solid unknown material in the cold trap used 
to condense any volatile iron compound. The amounts 
of iron measured as carbonyl were insignificant, but 
it is plausible to assume that a large fraction was 
decomposed either before it could leave the barrel or 
in the muzzle tube. So far, it has not been possible to 
identify these deposits, in part or in whole, with the 
decomposition products of iron pentacarbonyl. 

14 3 3 Vent Plugs 62 63 

Method 

The apparatus was identical in principle with that 
used to study erosion in vent plugs fired with propel¬ 
lants shown in Figure 5 of Chapter 11. The vent plugs 
were ^-in. long with a 34-in. bore diameter. They 
were made of SAE 4140 steel obtained from a liner 
for a 5-in./25-cal. gun. The bursting pressure of the 
rupture disk was 25,500 psi which corresponds to the 
maximum partial pressure of carbon monoxide and 
dioxide in guns. 

Weight Losses 

Figure 7 shows that the erosion weight losses are 
very strongly dependent on the flame temperature of 
the gases. Figure 8 shows that there is little if any 
dependence on the (CO + O 2 VCO 2 ratio e except 
when the C0 2 greatly exceeds the CO. f It is concluded 
that erosion under the conditions of these vent plug 

e For convenience we shall use the expression CO/CO 2 . The 
dissociation of carbon dioxide exceeds 2% only for flame tem¬ 
peratures above 2800 K but it also increases as the pressure 
drops. The presence of hydrogen lowers the amount of oxygen 
because of the water-gas reaction. 

f A ratio of CO/CO 2 < 1 in the gases from propellants is 
unlikely. It might occur with a composition greatly in excess 
of 40% of nitroglycerin. 107 


CONFIDENTIAL 







REACTION OF CARBON MONOXIDE WITH GUN STEEL 


291 


experiments was due to melting, with a possible ex¬ 
ception for the gases at 2160 K. 


X-Ray Analysis 

The material collected from the muzzle face of the 
vent plugs showed in all cases ferrous oxide (FeO), 
alpha-iron, austenite (gamma-iron containing dis¬ 
solved carbon), and in some cases magnetite (Fe 30 4 ). 
At 2160 K there was in addition hematite (Fe 2 03 ) 



Figure 7. The weight losses of %-in. vent plugs sub¬ 
jected to the action of eroding gases of controlled com¬ 
position and flame temperature resulting from the com¬ 
bustion of oxygen and carbon monoxide. (See Figure 
8.) (This figure appeared as Figure 9 in NDRC Re¬ 
port A-310.) 


present. The possibility that aerial oxidation occurred 
subsequent to the firing was excluded by other exper¬ 
iments. However, since the material was transported 
from the location of erosion to a position between the 
muzzle face of the vent plug and the abutting rupture 
disk, it is not necessarily representative of the actual 
reaction products of the gas with the gun steel. That 
the reaction of droplets of molten steel or primary 
reaction products, with the eroding gas, could well 
produce these erosion products was demonstrated by 


experiments with iron filings described in Section 
14.5.3. Microanalysis of the erosion products from 
the muzzle face indicated carbon in excess of the aus¬ 
tenite content and not present as a carbon compound 
in the x-ray analysis. The formation of cementite on 
the bore surface, later melted and dissociated in the 
melting and transportation, would yield free carbon 
and alpha-iron. g The investigators preferred to attri¬ 
bute the presence of apparently free carbon to the 
cracking of the carbon monoxide. The formation of 









< 

^34* 

\ 

10 K 





< 

\ 

h 

\ 

332C 

\ 

\ 

\ 

) K 

3170 

K 




w 

k \ 
s, \, 

l 

V 

\ 

V V 

\ 

■A 

\ 

> 2990 K 
\ 

\ 




\\ 

M 

sA 

\ 

-4 

\ 

^ 

\ ' 
N 

\ 

_ 

\ 

\ 

s v 

1^2765 K 
\ 

\ 

\ 

2495 K 

2160 K 


\ 

4 

_ 


O .58 .84 U5 1.45 1.89 2.5 

CO+Oj/COj 


Figure 8. The weight losses of the hrin. vent plugs 
in Figure 7 plotted as a function of the (CO + 0 2 ) /CO 2 
ration for various flame temperatures. (This figure ap¬ 
peared as Figure 8 in NDRC Report No. A-310.) 


cementite (Fe 3 C) does not necessarily require the 
presence of free carbon, 60 as shown by equations 
(5) and (10) of Chapter 13. 


Sulfur in the Charge 

Sulfur is present to the extent of 0.04 to 0.33% in 
all gun charges and it is converted almost entirely to 

* Apparently the solution of carbon in the liberated iron is 
too slow to stabilize the iron as austenite on cooling even in 
experiments where the decomposition occurred over a period 
of several seconds. 106 


CONFIDENTIAL 















































292 


EFFECTS OF CONSTITUENTS OF POWDER GASES 


hydrogen sulfide (H 2 S) in the powder gases, as ex¬ 
plained in Section 14.4.1. Catalysts for the formation 
of iron carbonyl mentioned in the chemical literature 
are organic and inorganic sulfides, ammonia and hy¬ 
drogen. 499 Hydrochloric acid and chlorine gas are 
reported to be inhibitors or destroyers of iron car¬ 
bonyl. 449 

In some of the vent plug experiments carbon di¬ 
sulfide, or sulfur dioxide or hydrogen sulfide was 
added to the mixtures of oxygen and carbon monox¬ 
ide having flame temperatures of 2330 and 2495 K. 
A continuous layer of iron sulfide was shown to be 
present on the bore surface of the vent plugs by sulfur 
prints, examples of which are shown in Figure 9. 
Carbon disulfide or sulfur dioxide did not essentially 
alter the erosion, but owing to experimental difficul¬ 
ties this result is not final. 


r 



Figure 9. Sulfur prints that indicate the formation of 
a layer of iron sulfide on the bore surface of vent plugs 
when: (top) carbon disulfide was added to the exploding 
gas mixture; and (bottom) 1.25% of hydrogen sulfide 
was added. (This figure appeared as part of Plate VII 
in NDRC Report No. A-310.) 

Hydrogen sulfide, even in amounts as small as a 
few tenths of 1 per cent, greatly increased the erosion, 
the increase being proportionately greater the lower 
the flame temperature and the higher the C0/C0 2 
ratio in the eroding gases. A maximum effect was 
reached at about 5% hydrogen sulfide. 

In a series of explosion mixtures that produced 
eroding gases of increasing flame temperatures, 1.25% 
of hydrogen sulfide was added. Whereas at the lower 
flame temperatures the relative effect of the hydrogen 
sulfide was large and dependent on the flame temper¬ 
ature, the increase in erosion caused by the addition 
of the hydrogen sulfide for flame temperatures from 
2500 to 3320 K was relatively small and practically 
constant. 

These results might be interpreted as indicating an 
erosion mechanism involving the formation of iron 
carbonyl particularly during the time that the bore 


surface is heating up, this reaction being catalyzed by 
the presence of hydrogen sulfide. Iron carbonyl was 
not isolated in vent plugs experiments, however, 
which may have been due to its inability to survive 
the peak temperatures in sufficient amount. 

On the other hand, the reaction 

Fe + H 2 S->FeS + H 2 

is highly exothermic; the products would be heated 
to between 600 and 700 C above the reactants in an 
adiabatic system. 60 It is worth noting that a similar 
reaction between iron and sulfur dioxide, although 
not known, should be much less exothermic if not 
endothermic. Any tendency for the ferrous sulfide 
(FeS) to be heated above the equilibrium tempera¬ 
ture of 1000 to 1100 C will result in its decomposition. 
If there is a stirring of the surface material, the melt¬ 
ing point could be lowered to about 985 C. Such an 
exothermic reaction could also contribute heat to af¬ 
fect the conditions at the bore surface in several ways. 
If bore-surface temperatures are just below a critical 
temperature, such as the melting point of erosion 
products, the formation of ferrous sulfide could con¬ 
tribute a small amount of heat that would be suffi¬ 
cient to cause this temperature to be reached, or, the 
rapidity of the temperature rise of the bore surface 
might be increased by the early formation of ferrous 
sulfide, thus prolonging the time that the bore surface 
is at or above some critical temperature. 

Both these effects should be particularly noticeable 
at low gas temperatures. At the higher flame tempera¬ 
tures, however, where melting is almost certainly the 
principal mechanism of erosion in 34-in. vent plugs, 
the exothermic heat would contribute in a relatively 
small and rather constant degree in the second man¬ 
ner suggested above. 

Some support to the foregoing hypothesis is offered 
by a few experiments with long vent plugs having a 
% 2 -i n - bore diameter. When the bore surface was 
coated with ferrous sulfide prior to the explosion, the 
weight losses were insignificant (1 and 4 mg) but 
when the surface was not so coated and instead 
1.25% of hydrogen sulfide was added to the exploding 
mixture the vent plug lost 43 mg. It should be em¬ 
phasized, however, that the data are too meager at 
present to substantiate either of the above points of 
view. 

X-ray analysis showed a trend to smaller amounts 
of FeS and larger amounts of FeO in the material 
from the muzzle face of the plugs as the flame tem¬ 
perature was increased and the CO/C0 2 ratio de- 


CONFIDENTIAL 








293 


REACTION OF CARBON MONOXIDE WITH GUN STEEL 


creased. This can probably be explained by the 
reaction: 

FeS + C0 2 - >FeO + CO + S 

Hydrogen sulfide is most effective when the bore 
surface is but mildly heated. This is illustrated by 
firings in vent plugs having >f 6 -in., J^-in., and %>-in. 
bore diameters which showed factors of increase of 
erosion caused by the addition of 1.25% of hydro¬ 
gen sulfide equal to 3, 10, and 66 respectively. 
This is fairly strong evidence that chemical factors 
are important. A simple explanation would be to 
assume the catalyzed formation of iron carbonyl, but 
it is not necessary, and in some cases it would be 
difficult, to postulate the formation of iron carbonyl 
to account for any appreciable portion of erosion in 
the absence of hydrogen sulfide. 

Ammonia and Hydrogen in the Charge 

The addition of 1 mole % of hydrogen sulfide, or 
ammonia, or hydrogen to the 2300 K mixture caused 
the weight losses to increase by more or less the same 
factor of about 3, from which one might infer that the 
increased erosion arises from the addition of 1, 1%, 
and 1 mole % of hydrogen respectively in the above 
gases. However, later work seems to rule out a com¬ 
mon mechanism for the increased erosion. The pres¬ 
ence of the above three gases also have a decided 
effect on the erosion of vent plugs of Armco iron, 
nickel, and cobalt in the same general manner. 

The addition of 1% of chlorine or hydrochloric acid 
had no effect on the Aveight losses, Avhich is remark¬ 
able in itself. That there was no mitigation of the 
erosion seems to demonstrate that iron carbonyl for¬ 
mation is insignificant in the absence of the above 
catalysts but there is no assurance that these inhibi¬ 
tors are acting in the explosion in the manner sug¬ 
gested. 

Extreme Severity of Tests 

Preliminary investigations Avere largely confined to 
the use of %-in. vent plugs rather than the in. size 
later studied. The conditions of erosion Avere ex¬ 
tremely severe Avith them, the chief part played by 
the gases being a transport of heat. The C0/C0 2 
ratios Avere all beloAv 2.5 Avith the one exception of a 
ratio of about 3. Studies of the composition of 
quenched poAA r der gases, described in Section 2.3, 
shoAv that these ratios are about the same as those 


obtained Avith double-base poAvders Avhich range from 
1.2 to 2.5. This is a condition which unfortunately 
cannot be remedied because <j)f the great difficulty of 
properly exploding such mixtures of carbon monoxide 
and oxygen that ^re necessary to give the higher 
C0/C0 2 ratios. 

Isolation and Identification of 
Iron Carbonyl 63 

The conditions for the synthesis of iron carbonyl 
are all satisfied in the passage of propellant gases 
through a gun barrel with the exception of the tem¬ 
perature. For a short period the temperature condi¬ 
tion is also satisfied near the origin of rifling and 
probably for a protracted period near the muzzle. 

Experimental Method 

The gases emerging from vent plugs 4- to 4J^-in. 
long Avith %2-in- bore diameter, or from a caliber .30 
barrel, were collected in an evacuated glass-lined 
chamber, hereinafter called the muzzle tube, shown in 
Figure 6 of Chapter 11. These gases were later drawn 
through a cold trap maintained at the temperature of 
liquid nitrogen. Any volatile matter condensed there¬ 
in Avas drawn through an analyzing system in which 
the pentacarbonyl would be decomposed by heat to 
form a thin mirror of iron-bearing material (analyzed 
for iron in quantitative Avork) and to form gaseous 
products Avhich were analyzed for the presence of 
carbon monoxide. 

The eroding gas mixtures Avere cool, having been 
obtained from an explosive mixture of 11 or 12% of 
oxygen in carbon monoxide. The use of large-bore 
vent plugs or caliber .30 barrels resulted in a much 
milder test than those obtained in the erosion tests 
of vent plugs just described in Section 14.3.3. The 
bursting pressure of the rupture disk Avas 25,500 psi. 
Caliber .30 barrels Avere fitted Avith a second 10,000- 
psi rupture disk at the muzzle. 

In order to favor the formation of iron carbonyl, 
up to 2.5% of hydrogen sulfide Avas added or the bore 
surface (of vent plugs only) was converted to ferrous 
sulfide before the test. A number of different arrange¬ 
ments in the‘muzzle tube Avere tried with the object 
of preserving any iron carbonyl formed. 

Location of Erosion Products 

Iron-bearing material was found at one or more of 


CONFIDENTIAL 






294 


EFFECTS OF CONSTITUENTS OF POWDER GASES 


three different positions. Firstly, a consistent feature 
was the appearance of an adherent iron-bearing de¬ 
posit on the wall of the muzzle tube in the form of a 
bright metallic mirror, reminiscent of the decompo¬ 
sition of a volatile metallic compound. This mirror 
became black if hydrogen sulfide was present in the 
eroding gases. Secondly, the cold trap usually con¬ 
tained a black sooty iron-bearing material. Thirdly, 
when iron carbonyl was collected a mirror of iron¬ 
bearing material was formed in the heated deposition 
tube of the analytical apparatus. Analyses for iron in 
the various deposits showed it to be present (1) as the 
element, (2) as an oxide, or (3) as a sulfide in cases 
where hydrogen sulfide or black powder had been 
used, but no sulfides of iron, or of copper often 
found in the deposits, were identified by x-ray 
analysis. 

Vent Plugs 

Muzzle Tube. X-ray examination of the mirror de¬ 
posit showed only ferrous oxide (FeO) and austenite. 
Under certain conditions ferrous oxide can be a prod¬ 
uct of the decomposition of iron pentacarbonyl but it 
could have been produced by the passage of hot ero¬ 
sion gases after the iron-bearing deposit was formed. 
It is, however, difficult to understand the presence of 
austenite by either of these mechanisms. Whether the 
deposit in the muzzle tube represents decomposition 
of iron carbonyl, and if so to what extent, requires 
further experimentation. 

Austenite and ferrous oxide were also obtained 
with a hydrogen-oxygen explosion but in this case the 
deposit was but loosely adherent and of different 
physical appearance. A comparison with the carbon 
monoxide-oxygen explosion is probably unfair pri¬ 
marily because of the very different thermal con¬ 
ditions. 

Cold Trap. The iron-bearing products in the cold 
trap could not be identified with the possible excep¬ 
tion of austenite. The x-ray patterns, except for a 
single line in one experiment, correspond neither to 
unidentified lines for the products of thermal decom¬ 
position of iron pentacarbonyl nor to lines found in a 
preparation of a condensed carbonyl, Fe 3 (CO)i 2 . 

Deposition Tube. None of the experiments with 
steel vent plugs revealed the presence of iron car¬ 
bonyl condensed in the cold trap. The possibility of 
its formation is not thereby ruled out, for iron car¬ 
bonyl introduced into the explosion mixture and fired 
through a copper vent plug failed to survive the ex¬ 


plosion. It was even lost to a marked extent on mere 
standing in the explosion mixture. Thus, it is quite 
likely that iron carbonyl may have been formed in 
the erosion of steel vent plugs in an amount less than 
the minimum for survival and detection. If mild ther¬ 
mal conditions were to be maintained the only solu¬ 
tion was to increase the area of the bore surface. This 
was accomplished by adapting caliber .30 Enfield 
barrels to the apparatus as shown in Figure 6 of 
Chapter 11. 

Gas Mixtures in Caliber .30 Barrels 

Bore-surface temperatures in the caliber .30 barrels 
were lower than in the vent plugs and a considerable 
area of the added surface will reach still lower maxi¬ 
mum temperatures. The correspondence between the 
temperature-time relations for this test and those for 
ordnance conditions with a projectile present has not 
been determined. 

Table 1 . Iron content of deposits after five consecutive 
firings through caliber .30 barrels used as vent plugs. 63 
(Propellant: carbon monoxide and oxygen mixture con¬ 
taining 11% oxygen and a little hydrogen sulfide.) 


Per cent 
H 2 S 

On glass 
liner of 
muzzle tube 

Milligrams of iron 

Solid 

material in 
cold trap 

From decom¬ 
position of 
condensed 
iron carbonyl 

0.00 

0.44 

3.02 

0.007 

0.125 

0.87 

3.52 

0.031 

0.25 

1.40 

5.72 

0.063 

0.50 

1.56 

8.53* 

0.046* 


* Some black material blown out of the cold trap and probably some iron 
carbonyl also lost. Blank runs on 4,000 psi of carbon monoxide without H 2 S 
and oxygen showed complete absence of iron carbonyl. 


No iron carbonyl was detected in a single firing but 
the usual mirror deposit was found in the muzzle 
tube. Iron carbonyl was repeatedly found in single 
firings if 1.25% of hydrogen sulfide was added. A 
series of firings were made in which small amounts of 
hydrogen sulfide were added. Analyses for iron in the 
three deposits yielded the results in Table 1. The 
addition of hydrogen sulfide, even in the small 
amounts present in propellant gases, increased the 
amount of iron recovered from the muzzle tube and 
cold trap and substantially increased the amount of 
volatile iron carbonyl recovered. 

Standard Propellant in Caliber .30 Barrels 
In order to produce the same number of moles of 


CONFIDENTIAL 







REACTION OF CARBON MONOXIDE WITH GUN STEEL 


295 


carbon monoxide 11 as given by the gas mixture, 20.5 g 
of an NH-M1 powder was burned in the same ex¬ 
plosion vessel with and without the addition of black 
powder. Piezoelectric measurements of the pressure 
as a function of time, illustrated in Figure 10, showed 



Figure 10. Oscillograph record of pressure as a func¬ 
tion of time for two types of firing in the same explosion 
vessel. (Upper part of figure): A typical experiment with 
the carbon monoxide and oxygen mixture. (Lower part 
of figure): One of the experiments with the NH-M1 
solid propellant. While the combustion of the gas mix¬ 
ture is comparatively very slow, the time of effusion 
through the barrel after rupture of the disk is the same 
for the gas mixture and solid propellant. (This figure 
appeared as Figure 33-1 in NDRC Report No. A-311.) 

that the maximum pressure, rate of decrease of pres¬ 
sure, and total time of effusion were about the same 
for the gas mixture and solid propellant. However, 
the very slow combustion of the gas mixture com¬ 
pared with the solid propellant is well demonstrated. 
In experiments with black powder (containing 10 per 
cent of sulfur) the amount added was the same as 
that normally present in the charges of a number of 
medium caliber guns. (See Section 14.4.1.) 

The deposits of iron-bearing material produced by 
the decomposition of the recovered iron carbonyl in 
the analytical apparatus are illustrated in Figure 11. 
Analyses for iron in the three deposits, given in Table 
2, show that iron carbonyl was formed with the NH- 
M1 propellant without the addition of sulfur in the 
form of black powder. The iron content of the deposit 
in the muzzle tube was increased by the addition of 
black powder, and most probably the iron present in 

h The amount of CO present in powder gas at any instant is, 
unlike the case of gas mixtures, very strongly dependent on the 
momentary temperature of the water-gas produced by the 
burning of the propellant. See the discussion in Section 2.3.4. 


Table 2. Iron content of deposits after five consecutive 
firings through caliber .30 barrels used as vent plugs. 63 
(Propellant: smokeless powder.*) 

— I 


Milligrams of iron 


Addition to 
powder 
(g) 

On glass 
liner of 
muzzle tube 

Solid 

material in 
cold trap 

From decom¬ 
position of 
condensed 
iron carbonyl 

0.0 

6.96 

38.64 

0.078 

0.42 black 
powder 

22.33 

21.30f 

0.058f 

1.0 p-nitro- 
benzoyl 
chloride 

16.31 

30.71 

0.089 


* Powder—NH-Ml for 37-mm gun, M1916, Lot 3727. 
t Material blown out of the cold trap very probably accompanied by loss 
of condensed iron carbonyl. Addition of black powder altered the nature of 
the explosion. 

the cold trap as iron carbonyl would have been 
greater had not the violence of the explosion caused 
a loss of material. The addition of paranitrobenzoyl 
chloride, added as a source of chlorine which has been 
reported to be an inhibitor for the formation of iron 
carbonyl, failed to reduce the amount of iron car¬ 
bonyl recovered and produced no change in the total 
iron from the three sources. 


Solid Iron-Bearing Deposits 

The amount of iron recovered as iron carbonyl 
from the gas mixture or the propellant, in the absence 



Figure 11 . Deposits of iron-bearing material from the 
thermal decomposition of iron carbonyl in the recovered 
gaseous products from firings of NH-Ml powder with 
and without the addition of different substances. (Up¬ 
per) NH-Ml propellant alone; (middle) 2% of black 
powder added; and (lower) 5% of paranitrobenzoyl 
chloride added. (This figure appeared as Plate XI in 
NDRC Report No. A-311.) 

of sulfur, was roughly constant at 0.2% of the total 
iron from all the analyzed deposits. The addition of 
hydrogen sulfide to the gas mixture increased the iron 
present as carbonyl to 0.8% of the total. These 


CONFIDENTIAL 


















296 


EFFECTS OF CONSTITUENTS OF POWDER GASES 


amounts are insignificant but it is plausible to assume 
that a large fraction of the iron carbonyl is decom¬ 
posed either before it can leave the barrel or in the 
muzzle tube. 

X-ray examination of the bore-surface products 
near the origin of rifling showed the presence of fer¬ 
rous oxide, alpha-iron, and sometimes austenite, with 
a few unidentified lines. For the firings of propellant 
plus black powder all the lines were attributable to 
alpha-iron and magnetite (Fe 3 0 4 ). The iron-bearing 
material in the muzzle tube and cold trap could not 
be identified; the unidentified pattern, except for one 
or possibly two lines, is not the same for the various 
deposits. It is curious that no known sulfides were 
identified, because their presence was obvious from 
the chemical analyses. 

The thermal decomposition of iron pentacarbonyl 
is strongly dependent upon the temperature and na¬ 
ture of the carrier gas. In a study of the factors 
affecting the decomposition one or more of the follow¬ 
ing products were found: alpha-iron, magnetite, 
cementite (Fe 3 C), always accompanied by unidenti¬ 
fied but often different lines. Two of these lines occur¬ 
red fairly consistently and were possibly the same as 
the pattern of the two unidentified lines that were the 
only lines found for the material from a muzzle tube 
after firing the gas mixture with hydrogen sulfide. 
Some of the unidentified lines in the decomposition 
products of iron pentacarbonyl corresponded to a 
part of the pattern obtained from a synthesized 
sample of the trimer of iron tetracarbonyl [Fe(CO) 4 ] 3 , 
but this latter pattern could not be repeated. 

It is unfortunate that the decomposition of iron 
pentacarbonyl has so far yielded no distinctive prod¬ 
uct, because the identified products are ones which 
might well be expected from reactions taking place in 
a gun tube quite independently of carbonyl forma¬ 
tion. It is thus too early to ascribe the presence of the 
solid deposits in part or in whole to the decomposition 
of iron pentacarbonyl. 

144 REACTION OF SULFUR GASES WITH 
GUN STEEL 

Introduction 

In the preceding section it was shown that the ad¬ 
dition of a small amount of hydrogen sulfide to the 
eroding gases produced by the combustion of carbon 
monoxide and oxygen resulted in a considerable in¬ 


crease in erosion. This section deals with the extent 
to which sulfur may form reaction products on the 
bore surface and the degree to which it penetrates the 
bore wall. If penetration is extensive, the physical 
properties of the steel can be affected adversely. If 
iron sulfide is a product of the interaction, the melting 
point of the bore surface material can be greatly 
depressed below that of steel. 

Elementary sulfur is present in the charges of 
medium and large bore guns to the extent of 0.04 to 
0.33% by weight of the charge, 53 by virtue of the 
booster and igniter pads of black powder. In small 
arms the sulfur, present to about the same extent, is 
all combined in the compounds of the primer. The 
sulfur goes almost completely to hydrogen sulfide in 
the powder gases according to a study of the chemical 
thermodynamics of gun erosion. 60 

The same study indicates that ferrous sulfide (FeS) 
should be the ultimate product under a great variety 
of conditions of temperature and gas composition for 
an interaction between a steel bore surface and pow¬ 
der gases derived from a propellant containing as little 
as 0.1% of sulfur. However, the formation of ferrous 
sulfide is self-inhibiting because of the highly exo¬ 
thermic reaction involved in its formation and this 
probably accounts for the lack of any considerable 
portion of ferrous sulfide identified in the reaction 
products at the bore surface of eroded guns described 
in Chapter 12. Another possibility is a scavenging of 
the sulfur by the usual deposit of copper in a gun 
tube, 98 which is discussed in Section 12.3.2. 

The extent to which the presence of black powder 
in a charge of smokeless powder contributes to ero¬ 
sion has been studied in an explosion vessel, but not 
under ordnance conditions. 53 ’ 200 - 203 

14-4-2 Summary of Results 

Sulfur in the reaction products on the bore surface 
and that which penetrated into the underlying steel 
was detected and measured by incorporating radio¬ 
active sulfur into a charge of single-base powder. This 
tracer method is sufficiently sensitive so that only a 
single firing is required and it possesses the advantage 
that it is often unnecessary to destroy the specimen 
lor a determination of the element in question. Thus 
it was possible to examine the sulfur content of the 
bore wall of a caliber .30 rifle barrel by using the 
whole barrel as a Geiger-Mueller counter in order to 
detect the emission of beta-rays from the radioactive 
sulfur. The same barrel could be studied in this man- 


CONFIDENTIAL 



REACTION OF SULFUR GASES WITH GUN STEEL 


297 


ner for the effect of a variety of treatments including 
the subsequent firing of a number of rounds. 

A considerable fraction of the sulfur in a single 
charge was found to be incorporated in the bore wall. 
Some of this sulfur may have been associated with 
the metal-fouling but evidence from various treat¬ 
ments of the barrel was in favor of an association of 
some of the sulfur with the steel, or an iron-bearing 
complex coating. Only a small fraction of the sulfur 
was permanently incorporated and 95% was lost in 
regularly decreasing amounts during the subsequent 
firing of 36 rounds. This loss is commensurate with 
the rate of erosion of caliber .30 barrels. The sulfur 
permanently incorporated would yield a maximum 
content of about 0.8% in the altered layer after 1,000 
rounds or alternately the crack system might contain 
material with a 2% sulfur content. 

This preliminary work led to further experiments 
which, while they departed from ordnance condi¬ 
tions, avoid ambiguous interpretations and provide 
a means of ascertaining the extent of penetration. 

Reaction products that were formed by the inter¬ 
action of propellant gases and the steel surface of test 
rods (described in Section 11.2.4) contained 12 to 
18% of sulfur by weight and this sulfur was shown to 
be combined with iron. The steel surfaces were chem¬ 
ically attacked by the sulfur components of the powder 
gases, and the formation of sulfur compounds of iron 
may contribute to erosion in guns. The evidence from 
these experiments, and others summarized below, 
appears to support the idea that the formation of a 
coating of reaction products in which iron from the 
steel is transferred to this coating is a necessary inter¬ 
mediate stage in erosion where melting of the steel 
itself is excluded, the chemical compounds so formed 
being more susceptible to complete removal than the 
original steel. 

The penetration of sulfur into the steel below the 
layer of reaction products is extremely shallow, al¬ 
though the sulfur content of this shallow zone is not 
inconsiderable. 

In view of these results a number of test rods were 
subjected to the action of powder gases derived from 
a nitrocellulose powder with and without the addition 
of black poAvder or its components. The addition of 
black poAA r der, and of sulfur or potassium nitrate in 
equivalent amount, increased the rate of erosion of 
gun steel as measured by weight loss. The addition of 
charcoal in equivalent amount decreased the rate of 
erosion and diminished the characteristic cracking of 
the surface commonly termed “heat checking.” While 


the addition of black powder always increased the 
erosion, curiously enough it greatly diminished the 
heat checking and even eliminated it A\dien in suffici¬ 
ent amount. There Avas abundant evidence that the 
gun steel per se was not melted in these tests and 
therefore the above observations could not be so 
explained. 

It Avas suggested on the basis of these experiments 
that an effort should be made to remove the sulfur 
from the charge in an actual firing test and perhaps to 
incorporate small amounts of charcoal that might 
mitigate erosion Avithout appreciably loAvering the 
potential of the poAvder. HoAvever, these things have 
not been accomplished. 1 

14 4 3 Experimental Methods 53 

Rifle Barrel as a Geiger-Mueller Counter 

In order to convert a rifle barrel to a Geiger-Mueller 
counter for detecting the radioactivity of a tracer ele¬ 
ment incorporated in the bore Avail it is only necessary 
to provide means for stretching a fine tungsten wire 
along the axis of the tube and to insure electrical 
insulation of this Avire and hermetic sealing of the 
tube. The location of the sulfur in the tube Avas de¬ 
termined by inserting glass tubes of appropriate 
lengths to block the beta-rays emitted by the radio¬ 
sulfur. Such an arrangement indicated maximum sul¬ 
fur content of the bore Avail over the first six calibers 
forAvard of the origin of rifling and a secondary max¬ 
imum at about 10 calibers behind the muzzle. The 
propellant was a standard IMR poAvder for caliber 
.30 small arms to Avhich 0.07% by Aveight of native 
sulfur Avas added. This sulfur acted as the carrier for 
the radiosulfur, the two being chemically indistin¬ 
guishable. The radiosulfur was shoAvn to index not 
only the added sulfur but all the sulfur in the charge, 
including that of the primer components. 

Penetration Experiments with Test Rods 

Small test rods were subjected to the action of 
straight nitrocellulose caliber .45 pistol poAvder gases 
in the apparatus described in Section 11.2.4 for test¬ 
ing the resistance of metals to surface cracking. Test 
rods were made of SAE 4140 steel in a tempered 

* The efficiency of certain ignition powders, including modi¬ 
fied black powders, that do not contain sulfur has been investi¬ 
gated at Aberdeen Proving Ground. Comparative erosion 
was studied only in vent plug tests. 200 ’ 203 


CONFIDENTIAL 




298 


EFFECTS OF CONSTITUENTS OF POWDER GASES 


martensitic state. The rods, % in- in diameter and 
134 in* long, were milled to give two diametrically 
opposite flat surfaces which were highly polished. 
These flats were subjected to the gases passing 
through two vents of D-shaped cross section formed 
by fitting the rods into a round hole in a replaceable 
cone. 

After less than the number of rounds that were 
established as sufficient to produce cracking had been 
fired, the rods were introduced into a specially de¬ 
signed Geiger-Mueller counter wherewith the sulfur 
content of the reaction products and the depth of 
penetration could be studied by the fall in activity 
following successive removals of known weights of 
material with a metallographer’s finishing lap. The 
amount of native sulfur which carried the radiosulfur 
was between 0.07 and 0.4% by weight of the smoke¬ 
less powder, corresponding to the amount present in 
some gun charges. 

Cracking and Erosion Tests with Test Rods 

The same apparatus, basic charge, and type of 
steel were used for these tests as for the penetration 
experiments just described. This kind of test posses¬ 
ses a decided advantage over the usual vent plug 
experiment in that it is possible to conduct a rapid 
microscopic examination of the surfaces at any stage 
before conclusion. The temperature-time relations 
are more readily controlled because only a portion of 
the gas passes over the specimen. The conditions 
were mild, as was confirmed by several observations 
that were made in order to establish the fact that the 
steel per se did not melt. Spectrophotometric meas¬ 
urements of the gas temperature of the sort described 
in Section 2.5.1 showed a maximum increase of 50 C 
over 2700 C with the addition of 10% of black pow¬ 
der. No change in the pressure-time curve was ob¬ 
served. A maximum weight loss of 12 mg after 85 
rounds was recorded with weighings on a micro- 
balance j or expressed in ordnance terms these tests 
correspond to an increase in bore diameter of 0.001 
to 0.005 in. in 30 to 60 rounds. This value was ob¬ 
tained by doubling the calculated thickness of metal 
removed, which is analogous to increase in groove 
depth. In some experiments black powder was added 
in amounts varying from 0.5 to 10% by weight; in 
other tests equivalent amounts of charcoal, sulfur, 
and potassium nitrate were added. 

j See Table 13 of Chapter 15 for weight losses of gun steel 
vent plugs of different sizes fired with different propellants. 


14,4,4 Sulfur in Reaction Products 53 

In all these experiments an adherent layer of ma¬ 
terial was produced on the surface that differed rad¬ 
ically from the original steel. The results of some 
experiments with test rods to determine the role of 
sulfur in the chemical alteration are given in Table 3. 
Each charge contained 0.4% of sulfur. 


Table 3. Sulfur content of reaction products on the 
surface of steel test rods subjected to the action of 
sulfur-containing powder gases. 53 


Number 
of rounds 

Thickness of 
coating 
(microns) 

Sulfur in coating 
determined by 
radioactive 
tracer (per cent) 

Computed 
as FeS 
(per cent) 

2 

0.8 

18 

50 

5 

1.8 

12 

33 

2* 

0.9 

12 

33 


* The charges were fired electrically in an all-steel system eliminating the 
possibility of copper, zinc, antimony, lead, or potassium that might have 
been derived from a caliber .30 case or its primer. 


The results in Table 3 do not prove the manner of 
occurrence of the sulfur. In other experiments no 
change in the activity of the surfaces was observed 
after removal of a loose deposit or after immersion in 
each of the following solvents in turn: ether, carbon 
disulfide, air-free water, and alcohol. These results 
indicated the absence of free sulfur in either of its 
modifications and established that the sulfur com¬ 
pound was not water-soluble or present in combina¬ 
tion with carbon and nitrogen in nonmetallic form. 

A Baumann or sulfur print affords an extremely 
sensitive method of detecting the presence of a sul- 
fide k of iron or manganese. The surfaces and cross 
sections of test rods fired with either free sulfur or 
black powder present in the charges gave very posi¬ 
tive sulfur prints similar to those illustrated in Figure 
9 for the cross section of a vent plug. 

An examination of the products on the surfaces of 
these test rods by electron diffraction produced a 
number of patterns which were difficult or impossible 
to identify, as explained in Section 11.5.2. Moreover, 
it is believed that some of the products yielded no 
patterns because they were amorphous. 98 With black 
powder or sulfur present in the charge the pattern 
corresponded to a double sulfide with a spinel struc¬ 
ture which was possibly violarite [(Ni,Fe) 3 S 4 ], al- 

k The use of this technique and a number of modifications 
devised to detect other sulfides, with particular reference to 
erosion products, is described in Section 11.5.3. 


CONFIDENTIAL 








REACTION OF SULFUR GASES WITH GUN STEEL 


299 


though it is likely that manganese or chromium could 
replace nickel in this compound. 137 There exist only 
meager data concerning such compounds. Thus, 
while the exact nature of the sulfur substance remains 
to be discovered, it has been demonstrated that iron¬ 
bearing compounds of sulfur are formed on steel sur¬ 
faces subjected to propellant gases containing only a 
small portion of sulfur. 

14 4 5 Penetration of Sulfur 53 

The penetration of sulfur into the steel beneath the 
coating of reaction products is shown in Table 4. 


Table 4. Penetration of sulfur into steel below the 
reaction products. 53 


Test 

rod 

Number 
of rounds 

Depth 

(microns) 

Sulfur content 
(per cent) 

5P 

2* 

0.00 to 0.26 

0.28 

6D 

2t 

0.00 to 0.19 

0.25 



0.19 to 0.26 

0.23 



0.26 to 0.57 

0.00 

6E 

5t 

0.00 to 0.16 

1.2 



0.16 to 0.24 

1.4 



0.24 to 0.31 

2.0 



0.31 to 0.39 

0.25 



0.39 to 0.66 

0.08 


* The rounds were fired electrically in an all-steel system. The sulfur 
content of the charge was 0.4% by weight of the smokeless powder fired. 

t The rounds were primer fired from a caliber .30 case which increased 
the total sulfur content of the charge to 0.5%. 

These values are averages for the whole affected area. 
Localized areas contained slightly more sulfur but 
not to greater depths, in similar experiments. The 
slight penetration does not follow the diffusion law. 
There appears to be a very thin zone of steel that has 
acquired a roughly constant amount of sulfur. An 
increase in the number of rounds increases the sulfur 
content without thickening this zone. The penetra¬ 
tion of sulfur is not too significant unless a very thin 
zone of “hot-short” material is formed which might 
aid the mechanical removal of the overlying reaction 
products. 

14 4 6 Cracking 53 

A number of test rods were studied with a compar¬ 
ison microscope in order to observe any surface ef¬ 
fects caused by the addition of black powder or its 
components to the charge of smokeless powder. The 
addition of up to 2% of black powder, i.e., less than 
the content of the charges for several guns, consider¬ 


ably diminishes the cracking of the surface. 1 * * This 
effect persists and is even more noticeable as the 
number of rounds is increased. With 6 or 10% of 
black powder, cracking is completely eliminated. 
With 10% of black powder the cracking is replaced 
by a pitting that resembles the ripple marks in sand. 

The black powder does more than prevent cracking 
of the surface in a given round because after test rods 
had been subjected to a number of rounds with 6 or 
10% of black powder in the charge they failed to 
exhibit cracking for a relatively large number of sub¬ 
sequent rounds without black powder present. After 
a survey of a number of possibilities it was concluded 
that the phenomenon is caused by a chemical change 
in the nature of the steel in the surface. This con¬ 
clusion agrees with the results of other studies dis¬ 
cussed in Section 13.5.3. The addition of sulfur alone 
intensified the cracking as shown in Figure 12. 
Potassium nitrate was even worse in this respect; but 
the addition of a third component in the form of 
powdered willow charcoal to the extent of 0.9% 
markedly reduced the cracking. 

144 * 7 Erosion 53 

The addition of black powder increases the rate of 
erosion. The results of weighing test rods are given in 
Figure 13. Similar curves were obtained for the addi¬ 
tion of sulfur and potassium nitrate equivalent to the 
6% black powder mixture but for the addition of the 
small equivalent amount of charcoal the erosion by 
comparison with pure smokeless powder was reduced 
by a factor of 2. The effects of the components of 
black powder in increasing the erosion caused by 
smokeless powder are not simply additive. There is 
fair agreement with the erosion test with black pow¬ 
der itself for a large number of rounds but a compar¬ 
ison of the initial erosion shows the very serious 
erosivity at this stage caused by potassium nitrate 
without its accompanying sulfur and charcoal. 

14 4 8 Charges without Sulfur 53 

The foregoing evidence obtained with test rods 
agrees with that obtained from vent plug tests with 
carbon monoxide-oxygen mixtures (Section 14.3.3) 
and from other vent plug tests 200 ’ 203 made at Aber¬ 
deen Proving Ground with propellant charges, in in- 

1 Such cracking is often termed “heat checking,” although 

this designation infers a cause that does not seem to have been 

proved. See Section 13.5.3. 


CONFIDENTIAL 









300 


EFFECTS OF CONSTITUENTS OF POWDER GASES 



Figure 12. The crack systems in the surface of steel 
test rods subjected to propellant gases, after 85 rounds. 
(Top): A nitrocellulose pistol powder. (Bottom): The 
same with 0.6% of native sulfur. Original magnifica¬ 
tion: 250X , reduced to about 220X in reproduction. 
(This figure has been taken from Figure 18 of NDRC 
Report No. A-276.) 

dicating the desirability of removing sulfur as an 
element in propellants in order to mitigate erosion, 
but only if the gun caliber and ballistic level are such 
that there is no melting of the steel per se. A study 
of the thermal effects of propellant gases in guns 48 
suggests that the latter condition is met for all stand¬ 
ard guns, with the exception of the 120-mm gun Ml 
and the possible exception of 16-in. guns. Therefore 
the removal of sulfur from the charge should mitigate 
erosion unless the temperatures are raised by high 
rates of fire. It would also seem desirable to avoid the 
presence of potassium nitrate, but it is possible that 
the correct proportion of the oxidizable component, 
charcoal, largely eliminates the effects caused by the 
presence of potassium nitrate alone. 


14 5 REACTION OF NITROGEN WITH GUN 
STEEL 

14 51 Early Investigations 

The role of nitrogen in the production of the fea¬ 
tures displayed on an eroded bore surface has been 
discussed for a number of years. 261 ’ 421 - 477 ’ 482 ’ 483 In 
these discussions the effect of solution of nitrogen in 
an austenitic layer was stressed rather than the oc¬ 
currence of definite nitrides of iron. Before the first 
paper on the nitrogen theory of erosion appeared, 
however, it had been suggested 478 that mechanical 
removal of a brittle nitride layer formed by nitrogen- 
ization of the bore surface might be an important 
mechanism in producing gun erosion. 

Nitrogen determinations of layers machined from 
the bore surfaces of badly eroded caliber .30 machine 
gun barrels showed an appreciable increase in nitro¬ 
gen content near the bore surface to a depth of 
several thousandths of an inch. 481 

14 5 2 Summary of Recent Results 

The thermodynamical instability of the iron ni¬ 
trides at high temperatures would seem to preclude 
their formation and preservation in a gun tube. 60 
Nevertheless, nitrides have been identified in worn 
Service guns (Chapter 12). Moreover it was shown 
that hexagonal (epsilon phase) iron nitride (Fe 2 N x ) 
constituted no small portion of the products of inter¬ 
action between gun steel filings and the gases from a 



Figure 13. The total weight loss of steel from test rods 
subjected to the action of propellant gases from a nitro¬ 
cellulose pistol powder to which different amounts of 
black powder were added. Curve (A) Pistol powder with¬ 
out added black powder; Curve (B) 6% of black powder 
added; Curve (C) 10% of black powder added. (This 
figure appeared as Figure 17 in NDRC Report No. 
A-276.) 


CONFIDENTIAL 




REACTION OF NITROGEN WITH GUN STEEL 


301 


Table 5. X-ray identification of the products of reaction between powder gases and pulverized materials. The numbers 
indicate only the order of abundance; 1 is the most abundant. Letter d: diminished amount with respect to preceding 
column; similarly, sd: slightly diminished; and gd: greatly diminished. 


Materials fired with 
I MR powder (single-base) 
Electrolytic iron Gun steel Fe 4 N 

325 200 100 200 

Product (Mesh) (Mesh) (Mesh) (Mesh) 


Fe 2 N* 

3 

2d 

3 sd 

2 

1 

Fe 4 N 





5 

FeO 

2 

2 gd 

3 sd 

2 

3 

Austenite § 

1 

1 

2 

1 

2 

Ferrite § 

4 

1 

1 

1 

4 


Fe 2 N 


1 * 


Materials fired with 
FNH-M2 (double-base) powder 
Iron Fe 4 N Fe 2 N x 

325 

(Mesh) 


2 2 * 


2 If It If 

13 3 3 

3 2 3 


* Lower nitrogen content than original. t 90 per cent. J Mostly FeO. § See footnote of Chapter 12 on page 248. 


single-base powder fired in the conventional manner 
from a caliber .30 rifle. 28 With the hotter FNH-M2 
powder no iron nitrides were formed. It appears that 
the nitrides must be formed either very late in a firing 
or as a secondary product. The main difficulty in 
establishing the mechanism of their formation is the 
known inertness of molecular nitrogen. Furthermore, 
ammonia, which is an active nitriding agent, is pres¬ 
ent in powder gases in only very small amounts. It 
is therefore interesting to find in other experiments 70 
that while an increase in total nitrogen is confined to 
the immediate surface, an exchange between the ni¬ 
trogen of the propellant gases and the original nitro¬ 
gen of the steel can occur to a very appreciable depth. 

14 5 3 Formation of Nitrides 28 

In order to expedite an x-ray study of the erosion 
products in a gun a very convenient and rapid ex¬ 
perimental procedure was devised as described in 
Section 11.2.6. This consisted of incorporating the 
pertinent substance in pulverized form in the powder 
charge for an otherwise normal round of a caliber .30 
rifle. The rifle was fired into an evacuated tube from 
which the products of the explosion were later re¬ 
moved for examination. Such an experiment conforms 
to ordnance conditions except for one important de¬ 
parture. The temperature of the finely divided ma¬ 
terial will rise much higher than that of the bore 
surface which is being simulated by the pulverized 
substance. The finer particles melted on firing; coarser 
particles were rounded at the edges only. The results 
are summarized in Table 5. ra 

m Compiled from a description of the results in the NDRC 
report 28 on this investigation, in which the relative amounts of 
the products are discussed in greater detail. 


Ferrous oxide (FeO) and austenite 11 were common 
to all the experiments. Ferrous oxide was by far the 
preponderant product with a nitroglycerin double¬ 
base powder. Hexagonal iron nitride formed a lesser, 
though still appreciable, portion of the total reaction 
product with IMR powder but was absent with the 
hotter FNH-M2 powder. This latter relation was 
fully confirmed in later work on the x-ray and chem¬ 
ical examination of eroded bore surfaces, described in 
Chapter 12. The conclusion to be drawn from the 
firings with the iron nitrides is that they are unstable 
under the temperature and pressure conditions of the 
experiments. The hexagonal iron nitride found in all 
the firings of iron and steel particles with single-base 
powder was therefore presumably formed as a second¬ 
ary product. 

14 5 4 Mode of Formation of Nitrides 

The mechanism of the formation of iron nitride on 
a gun bore surface is far from clear. Iron does not 
react with molecular nitrogen even at elevated tem¬ 
peratures and pressures; although at a partial pressure 
of 350 atm, which corresponds roughly to the maxi¬ 
mum powder-gas pressure in a gun, the solubility of 
molecular nitrogen in gamma iron may rise to around 
0.4%.° Upon cooling and release of pressure some of 
this nitrogen may be released and form nitrides. 

Ammonia reacts rapidly with iron to form nitrides, 
the reaction proceeding by release of monatomic ni¬ 
trogen at the surface. The amount of ammonia pres¬ 
ent in the gases from an FNH-M1 powder has been 

n See footnote of Chapter 12 on page 248. 

° There are no experimental data at such pressures. This 
value was calculated on the basis that the solubility is propor¬ 
tional to the square root of the pressure. 


CONFIDENTIAL 












302 


EFFECTS OF CONSTITUENTS OF POWDER GASES 


calculated to be roughly 0.2 mole % at a gas temper¬ 
ature of 1800 K and a gas density of 0.3 g per cc, 
falling to 0.06% at 1000 K and 0.1 g per cc. 60 Gases 
of these compositions can be in contact with bore 
surfaces at much lower temperatures because imme¬ 
diately next to the bore wall the gases, although at 
the bore surface temperature, probably have more 
nearly the equilibrium composition of the bulk of the 
gas in the tube. These low ammonia concentrations 
appear to be unfavorable to the formation of nitrides, 
particularly if the stability of the nitrides is such that 
one must assume them to be produced late in the 
firing. However, the nascent hydrogen released at the 
surface when any ammonia present dissociates will 
favor a replenishment of the adsorbed layer of am¬ 
monia by reacting with molecular nitrogen. The sol¬ 
ubility of nitrogen introduced into gamma iron in 
this manner need not be limited to the amount @f 
molecular nitrogen that might dissolve. 

Well-spheroidized carbides present in an alloy steel 
proved upon segregation to contain the total nitrogen 
of the steel, nitrogen being present to the extent of 
1 atom to every 23 of carbon. Because it was not 
possible from x-ray identification to detect the pres¬ 
ence of a discrete nitride phase, nitrogen may have 
been present as replacement or “floating” atoms. 98 
The available nitrogen being limited, this experiment 
does not establish the maximum extent of association 
of nitrogen with the carbides. This capacity of the 
carbides to scavenge the nitrogen may be important 
in reactions at the bore surface, where cementite is a 
commonly occurring product. The nitrogen thus con¬ 
centrated, if in sufficient amount, may separate as a 
discrete nitride phase on cooling, or possibly it may 
merely change the stability relations of the iron ni¬ 
trides at the high temperatures reached by gun-bore 
surfaces. It is generally true that the iron nitrides 
appear in conjunction with carbide in the reaction 
products of an eroded bore surface. 98 

While it is difficult with our present knowledge to 
understand the reason for the formation of nitrides 
of iron under the temperature conditions encountered 
during the firing of a gun, the experiments described 
below indicate a high mobility of original and ac¬ 
quired nitrogen well below the bore surface combined 
with enrichment close to the surface. 

14 5 5 Experimental Determination of 
Nitrogen Penetration 70 

Studies of badly eroded guns, as mentioned in Sec¬ 


tion 14.5.1, showed an increase in nitrogen content 
near the bore surface. It has now been shown in the 
manner described below that this enrichment is not 
solely associated with a badly eroded surface but is a 
feature accompanying the first firings through a new 
barrel. These experiments were under ordnance con¬ 
ditions. 

Method 

In order to detect the penetration of nitrogen ac¬ 
quired from the propellant gases into the bore wall, 
the nitrogen atoms of the propellant were tagged by 
changing the ratio of the two stable isotopes of nitro¬ 
gen having masses 14 and 15 P . This was accomplished 
by adding to a standard I MR powder for the caliber 
.30 rifle a small amount of ammonium nitrate that 
had been enriched in the rarer isotope of mass 15. 

A caliber .30 Springfield rifle, M1903A1 was assem¬ 
bled with an unproofed barrel and ten rounds were 
fired. Each round was standard in all respects except 
for the addition of 20 mg of the ammonium nitrate 
enriched with respect to nitrogen 15. The barrels were 
suitably cleaned and samples were taken by drilling 
along the axis of the bore. These samples were ana¬ 
lyzed for total nitrogen content, following which the 
residue of ammonium sulfate was treated to recover 
the nitrogen as the element. 

The atomic percentage of nitrogen 15 was then 
determined in these samples of nitrogen with a mass 
spectrometer. Ordinary nitrogen (for example, the 
original nitrogen contained in the steel) has 0.372% 
of nitrogen 15. The nitrogen in the propellant gases 
contained, by virtue of the addition of the ammonium 
nitrate, 0.642% of nitrogen 15. From these values 
and the atomic analysis of the nitrogen in a given 
sample it was possible to calculate that fraction of 
the total nitrogen in the sample which is to be attri¬ 
buted to the presence of propellant nitrogen. 

Validity of Assumptions 

No experiment was performed to demonstrate defi¬ 
nitely a complete dissociation of the ammonium ni¬ 
trate but it is safe to assume complete dissociation in 


p The use of a radioactive isotope of nitrogen in the manner 
described in Section 14.2.3 for tracing carbon would have 
greatly aided an exploration of very thin layers. The half-life 
of radio nitrogen (io min) precluded its use in this type of 
experiment; also it would have failed to show the features of 
exchange without enrichment. 


CONFIDENTIAL 





REACTION OF NITROGEN WITH GUN STEEL 


303 


view of the experience with barium carbonate de¬ 
scribed in Section 14.2.4. It was established that the 
ballistic level was unaffected by the addition of 0.6% 
by weight of ammonium nitrate in 3.24 g of the I MR 
powder. 


Limitations on Interpretations of Results 

The results must be interpreted in the light of the 
following limitations. 

1. Measurements were made on surfaces which ex¬ 
tended over one-third to one-half of the length of a 
barrel. Consequently areal differences in nitrogen 
content and exchange may be averaged out to a large 
extent. 

2. The depth of any given cut was at least 0.0025 
in. (64/x) and therefore a feature of the penetration 
occurring in a very thin zone would have been largely 
masked. This is unfortunate with respect to the in¬ 
crease in total nitrogen content of the surface cut, 
but is unimportant in discussing features of nitrogen 
exchange. 

14,5 6 Bore-Surface Temperatures 70 

No attempt was made to simulate bore-surface 
temperatures occurring in larger bore guns, as de¬ 
scribed in Section 14.2.4 for the work on carbon. 
The calculated maximum peak temperature of the 
bore surface was therefore around 630 C" for barrel 
C. Barrel B was fired at the rate of about 1 round per 
second, which would raise this value for the last 
round by an estimated 60 C. These temperatures are 
somewhat in excess of the eutectoidal temperature of 
the iron-nitrogen system, which is about 600 C. 

14,5,7 Nitrogen Penetration 70 

The results of the nitrogen penetration determina¬ 
tions are shown in Figure 14. Barrel B was a six-land 
barrel made of WD 1350 steel containing 1.5% of 
manganese but no chromium or molybdenum. Barrel 
C was a four-land barrel made of WD 4150 steel. The 
following points are to be noted. 

1. Increase in total nitrogen content is confined to 
the surface cuts (the first and second bars of the 
graphs where lands and grooves were separated). 

2. Nitrogen from the propellant gases penetrates to 
a comparatively great depth but merely replaces 
some of the original nitrogen with no change in the 
total nitrogen content. 



Figure 14. The penetration of nitrogen from propel¬ 
lant gases derived from IMR powder into the bore wall 
of unproofed caliber .30 steel rifle barrels. The bars 
indicate total nitrogen content; the black areas show, 
with a magnification of 5X, that portion of the total to 
be attributed to nitrogen acquired from the propellant 
gases. Bb, breech third of barrel B, Be center third pf 
barrel B, Bm muzzle third of barrel B, Cb breech half 
of barrel C, Cm muzzle half of barrel C. (This figure 
was based on Figures 2 and 3 in NDRC Report No. 
A-398.) 


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304 


EFFECTS OF CONSTITUENTS OF POWDER GASES 


3. Nitrogen from the propellant gases is insufficient 
to account for the increase in total nitrogen in the 
surface. It appears that displaced nitrogen below the 
surface eventually travels counter to the replacing 
nitrogen and is finally located in the surface. This 
would provide a mechanism for the late formation of 
nitrides. In the light of present knowledge, no other 
explanation can be offered for this increase. 

4. In general, the proportion of nitrogen from the 
propellant gases is highest in the surface cut but there 
is a secondary maximum at the surprising depth of 
0.015 to 0.025 in. (measured from the groove sur¬ 
faces) without any measurable difference in the total 
nitrogen content. 

5. The depth of the secondary maximum of nitrogen 
exchange is roughly the same irrespective of the loca¬ 
tion in the barrel. It occurs somewhat closer to the 
bore surface in barrel B subjected to rapid fire and is 
absent in the muzzle third. However, this barrel had 
a different composition from barrel C. 

6. The total depth of nitrogen exchange is greatest 
in the breech third of barrel B. This depth differs but 
little for the two halves of barrel C. The total depth 
of nitrogen exchange is greatest, near the breech, in 
the barrel subjected to rapid fire; but again this dif¬ 
ference may be due to the different composition of 
the steel. 

Nitrogen Exchange 

The high degree of mobility of nitrogen at rela¬ 
tively low temperatures is well illustrated by these 
experiments, and yet it appears that there is a strong 
controlling influence which, while it permits the ex¬ 
change of nitrogen at depth, precludes any measur¬ 
able addition of nitrogen except to a layer right at 
the bore surface. Such an influence might be attrib¬ 
uted to the carbides which would be present in 
fairly constant amount except right at the surface 
where a considerable increase in carbide content is to 
be expected on the evidence in Chapter 12 and in 
Section 14.2. If it is true that in these steels the nitro¬ 
gen is completely associated with the carbides as was 
found to be the case in a well-spheroidized alloy steel, 98 
then exchange of nitrogen is associated with the car¬ 
bide grains. 

The subsurface maximum of propellant nitrogen is 
unexplained. At a depth of 0.02 in. the maximum 
temperature near the origin of rifling during a single 
firing in a caliber .50 barrel is about 100 C and about 
50 C near the muzzle. 106 In a caliber .30 barrel these 


temperatures will be lower. It therefore appears that 
some other variable than temperature may be more 
important in controlling penetration and exchange, 
perhaps the pressure of the powder gases or the mo¬ 
mentary stress configurations set up in the gun tube 
during firing. (Section 7.1.) The remarkable similar¬ 
ity of the nitrogen exchange over considerable lengths 
of barrel lends support to this argument. 

Nitrogen Increase 

It can hardly be supposed that the surface increase 
in nitrogen is distributed uniformly throughout the 
surface cut of arbitrary thickness. It seems reasonable 
to assume that the increase in total nitrogen is largely 
confined to a layer of the reaction products. In the 
study of carbon penetration the thickness of this 
layer could be determined. It was roughly 1 n thick 
near the origin of rifling for a barrel initially at 27 C." 
Table 6 indicates that the nitrogen content in such a 


Table 6. Nitrogen content of a layer of reaction prod¬ 
ucts 1 micron thick if the increase in nitrogen content 
of the surface cut is confined to such a layer. 


Barrel 

and 

location 

Increase in total 
nitrogen content 
of the cut 
(per cent) 

Nitrogen in a 
surface layer 

1 micron thick* 
(per cent) 

Barrel B Lands 

0.0029 

0.26 

Breech third Grooves 

0.0016 

0.16 

Barrel C Lands 

0.0040 

0.37 

Breech half Grooves 

0.0009 

0.08 


* Density assumed to be the same as that of the steel. 


layer would not be unreasonably large and that the 
concentration on the lands would be higher than in 
the grooves. In analyses of surface flakes removed 
from the eroded bore surface of a 3-in. gun by 
the method described in Section 11.4.4, the flakes 
from the lands contained 1.1% nitrogen while the 
flakes from the grooves showed only 0.3% nitrogen, 98 
which yields a ratio roughly the same as for the hypo¬ 
thetical calculation in Table 6 for barrel C. It appears 
that nitride formation is more severe on the lands of 
a gun than in the grooves. A further discussion of the 
analyses of the flakes mentioned above may be found 
in Section 12.5.2. 

Metallographic observations of cross sections of 
barrel B showed no development of cracks and 
no discernible change of structure caused by the 
firing. 


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REACTION OF THE INDIVIDUAL POWDER GAS CONSTITUENTS 


305 


14 6 REACTION OF THE INDIVIDUAL 
POWDER GAS CONSTITUENTS WITH 
GUN STEEL 

1461 Introduction 

The erosive action of a single constituent of the 
powder gas cannot be studied by attaining the con¬ 
ditions of temperature and pressure directly from the 
chemical energy of an explosion. An ideal method 
consists of compressing the gas adiabatically while it 
is confined by a rupture disk which breaks and re¬ 
leases the gas through the specimen after the neces¬ 
sary temperature and pressure have been reached. 101 
This constitutes a radical departure from ordnance 
conditions, but it is only by experiments of this type 
that one can appraise the erosive action of an individ¬ 
ual constituent and follow the changes of such action 
with the addition of the other constituents in mix¬ 
tures of progressively increasing complexity. Further, 
it should be possible to separate purely thermal effects 
from chemical effects by the use of an inert gas such 
as argon. 

14-6,2 Summary of Results 

A study of erosion and the nature of reaction 
products that result when appropriate single gases 
and combinations of these gases are heated by adia¬ 
batic compression and passed over the surfaces of gun 
steel did not provide much information. This was 
almost certainly due to the limitations of the gas 
capacity and energy of the available apparatus. 
Many gases and mixtures were used but only oxygen 
produced a measurable erosion. Of the several known 
reaction products (described in Chapter 12) that are 
found on the bore surface of a gun, only magnetite 
could be identified by electron diffraction 137 as a 
product of the gas-steel interaction in these experi¬ 
ments. Other compounds were either debris or un¬ 
known substances. 

Considerable thought was given to the fundamen¬ 
tal requirements and to the design of an apparatus to 
overcome the limitations of gas capacity and energy. 
The work was discontinued because of more pressing 
investigations. A continuation along the lines sug¬ 
gested here should yield valuable information. 

14-6-3 Experimental Method 101 

The apparatus for adiabatic compression experi¬ 
ments has been described in Section 11.3.1. Speci¬ 


mens were of two types and were made of SAE 4140 
steel obtained from a 5-in./25-cal. gun. For the deter¬ 
mination of erosion by weight loss, disks were used 
which wereO.06 in. thickwith center holes, 0.014,0.022, 
or 0.031 in. in diameter. For examination by visual, 
x-ray, or electron diffraction methods the specimen 
consisted of two small rectangular blocks, with the 
pertinent surfaces polished, which were clamped in a 
holder so as to form a slit about 0.010 in. wide. It is 
important to know the pressure-time relation which, 
apart from interest in this relation itself, yields the 
data to calculate gas temperature. A novel form of 
piezoelectric gauge was incorporated in the appara¬ 
tus. Experiments were conducted with helium, argon, 
nitrogen, hydrogen, oxygen, carbon monoxide, car¬ 
bon dioxide, air, and such mixtures as are indicated 
in Table 7. 


Table 7. Maximum temperature and pressure of 
adiabatically compressed gases in a series of tests for 
the erosive action on gun steel. 101 


Gas 

composition 
(vol %) 

Maximum 

pressure* 

(atm) 

Maximum 

temperature* 

(C) 

99£N2, IO 2 

800-1150 

1600-1750 

Argon 

750-1030 

3400-3800 

Hydrogen 

1200-1800 

1750-2000 

99£N 2 , IH 2 

900-1100 

1650-1750 

Hydrogen 

1300-3000 

1700-2100 

98 CO, 20 a 

1000-1600 

1550-1800 

A mixture! 

1300-2600 

1250-1500 

99£CO, |H 2 S 

650-1850 

1600-2100 


* The two values indicate the limit of scatter in 10 shots, 
t 36 CO 2 , 36 H 2 , 17 CO, 11 N 2 , which is a mixture with approximately 
the same atomic composition as nitrocellulose. 


14 6 4 Temperature and Pressure 

The ratio of the final temperature T 2 of the gas to 
its initial temperature Ti is a function of the ratio of 
the final pressure P 2 to the initial pressure Pi as ex¬ 
pressed by equation (1), 

T* (Pt\ « m 

Ti \Pi/ C v + R’ w 

in which R is the gas constant and C v the specific 
heat at constant volume. The final gas pressure is 
theoretically proportional to the initial pressure for 
a given gas and a fixed amount of work. 

Unfortunately, a difficulty with the available ap¬ 
paratus was a bouncing of the piston and falling 
weight combination that is well illustrated in a rep¬ 
resentative pressure-time record in Figure 15. The 


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306 


EFFECTS OF CONSTITUENTS OF POWDER GASES 


gas was sometimes compressed and re-expanded sev¬ 
eral times, with the result that a major portion of it 
escaped while the pressure and temperature were low. 

Table 7 gives the maximum pressures and temper¬ 
atures attained by various gases and mixtures and 
indicates the variation from shot to shot. When the 
maximum recorded pressures in Table 7 are compared 
with those of the powder gases, 107 it is found that 
these pressures exceed the partial pressure of nitrogen 
and hydrogen in a gun but are about correct for car¬ 
bon monoxide. The total pressure is somewhat low 
for the mixture simulating the composition of nitro¬ 
cellulose. It is important to note, however, from 
Figure 15 that, whereas the total time of efflux was 



Figure 15. Oscillograph record of the pressure as a 
function of time in the apparatus for compressing gases 
adiabatically. The dotted line is the 500-atmosphere 
ordinate. The time signals on the abscissa are 1 msec 
apart. (This figure appeared as Figure 6 in NDRC 
Report No. A-429.) 

of the desired order of magnitude, these ordnance 
pressures were sustained for only a fraction of the 
total time and that the specimen was alternately ex¬ 
posed to cooled gas and gas at the desired temper¬ 
ature and pressure. 

It is more difficult to evaluate the appropriateness 
of the gas temperatures. Spectrophotometric meas¬ 
urements of the gas temperatures in a gun reported 
in Section 2.5.5 would indicate that the maximum 
temperatures in Table 7, with the exception of argon, 
are somewhat lower than desired, particularly for the 
mixture. 

14 6 5 Erosion 101 

Air, helium, nitrogen, hydrogen, carbon dioxide, 
and carbon monoxide failed to produce a measurable 


weight loss or increase in bore diameter of the speci¬ 
mens. Oxygen alone gave an appreciable effect. In 
three shots the bore enlarged from 0.014 to 0.15 in. 

14 6 6 Reaction Products 

The surfaces of the blocks subjected to the action 
of the gases in the slit were studied by the methods 
of x-ray and electron diffraction (Section 11.5.2). The 
action of nitrogen indicated no change, but with air 
x-ray patterns of austenite, magnetite (Fe 3 0 4 ), and 
a trace of ferrous oxide (FeO) were obtained in addi¬ 
tion to those of the steel. 

Some 18 specimens were studied by electron dif¬ 
fraction 137 including those subjected to the gas mix¬ 
tures in Table 7. The only product which was prob¬ 
ably formed by a gas-steel interaction was magnetite, 
which was sometimes identified, particularly when 
small amounts of oxygen were added to the gases. In 
a study of the action of propellant gases on gun steel, 
a very thin film of magnetite was often found cover¬ 
ing other reaction products, 31 but although the me¬ 
thods devised to expose these subsurface reaction 
products to the electron beam were applied in detail 
to the present specimens, no underlying products 
were identified. 

On many of the specimens subjected to multiple 
firing, 5 to 10 shots, there was evidence of debris in 
the form of degenerate graphitic carbon, loose mag¬ 
netite, and organic compounds that resulted from a 
deterioration of the packing. This undesirable feature 
was eliminated by firing a single shot. 

In many cases part of the pattern was not iden¬ 
tified, but this unknown pattern was only identical for 
three specimens subjected to nitrogen. Of the known 
substances that result from the interaction of propel¬ 
lant gases with gun steel such as cementite (Fe 3 C), 
iron nitrides (Fe 4 N and Fe 2 N x ), complex iron cya¬ 
nides, wiistite or ferrous oxide (FeO), austenite (gam¬ 
ma-iron containing dissolved C or N or both) and 
magnetite (Fe 3 0 4 ), only magnetite was identified. 

The requirements necessary to duplicate the sever¬ 
ity of the usual vent-plug test for erosion are outlined 
below. The tests performed were only about one- 
tenth as severe and it is apparent that the conditions 
of these tests were too mild. 

14,6,7 Improvements in Apparatus 101 

As a result of the experience gained with the avail¬ 
able apparatus and a study of its shortcomings indi- 


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MODIFICATIONS OF PROPELLANTS 


307 


cated in the preceding sections, the experimenters 
were able to outline the following requirements for 
an adequate apparatus and to suggest a possible 
design of this apparatus. 

1. A sixfold increase in the mass of gas. 

2. An increase in the energy per unit mass of gas by 
a factor of 1.5. 

3. A means of arresting the compressing mechan¬ 
ism so that the gas is not alternately compressed and 
re-expanded. 

One solution would be to increase the dimensions 
of the already bulky apparatus by a factor of about 
1.8 but a major difficulty would arise in providing 
for requirement (3) above. 

An apparatus which was to have been built is 
shown schematically in Figure 16. Energy is derived 
from a solid propellant burning in the cylinder on the 
left. The gas is compressed in a single-stroke two- 
stage compressor on the right. The large-bore piston 
head shears off at the end of the low-pressure stage 
and permits the small-bore piston to complete the 
stroke in the high pressure cylinder of much smaller 
diameter. The piston arrester consists of two wedging 
blocks that are tripped by the passage of the head of 
the driving piston and thereupon close together be¬ 


hind this head. The thrust that these blocks must 
withstand is materially lessened by introducing the 
small-bore cylinder for the final compression. Such an 
apparatus should be extremely useful in studying the 
thermal or chemical effects of gases at high tempera¬ 
ture and pressure on a variety of materials when it is 
desirable to choose a specific gas or a combination of 
gases other than the qualitatively invariant mixture 
from a propellant. 


14 7 MODIFICATIONS OF PROPELLANTS 

Any radical modification of the propellant that 
might suppress one or more of the undesirable re¬ 
actions of the constituents of the powder gases with 
gun steel would require a very extensive research. A 
few experiment^, described in Section 15.6, were per¬ 
formed to test the possibilities of a slight but prac¬ 
tical change in the powder. It is now possible to state 
as a broad principle that double-base powders, which 
have high flame temperatures and give gases rich in 
carbon dioxide, cause oxidation of the bore surface, 
while the cooler single-base powders, which produce 
more carbon monoxide, are largely carburizing. 



Figure 16. Sketch of an apparatus designed to compress gases adiabatically which should perform without the un¬ 
desirable features inherent in the available compressor. (This figure appeared as Figure 7 in NDRC Report No. A-429.) 


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Chapter 15 


EROSION OF GUN STEEL BY DIFFERENT PROPELLANTS a 

By John N. Hobstetter h 


is 1 INTRODUCTION 

T he testing of gun steel with different propellant 
powders was undertaken by Division 1, NDRC, 
with a triple aim: to develop accelerated methods for 
testing and rapid classification of powders according 
to their erosiveness, to classify many powders on a 
scale of relative erosiveness, and to discover relation¬ 
ships between observed erosive effects and funda¬ 
mental properties of the powders. These three aims 
were amply realized in the course of the extended 
program of testing. 

A survey of the literature 16 had shown that the 
dependence of the magnitude of erosion on the type 
of propellant had been suspected as soon as nitrocel¬ 
lulose powders were introduced. Experimental firing 
had proved the point that not only was double-base 
powder more erosive than single-base powder but the 
amount of erosion with each type of powder depended 
on the barrel temperature. Vent-plug tests, in which 
erosion was predominantly by melting (as was the 
case also in the experiments described in Section 
14.3.2) had shown a rough correlation between ero¬ 
sion and the flame temperatures of the powders. 
Division 1 was able, by conducting an extensive series 
of accelerated tests under carefully controlled con¬ 
ditions of different degrees of severity, to determine 
quantitative relationships among types of powders, 
their properties, and the type and magnitude of ero¬ 
sion for many different powders. 

Two different approaches were made to the prob¬ 
lem of accelerated testing methods. The most direct 
one was the actual firing of gun steel barrels with the 
different propellants under conditions of high veloc¬ 
ity (3,300 fps) and of hypervelocity (3,600 fps). The 
less direct approach simulated firing conditions in a 
gun (except for the effects produced by a projectile) 
by means of a vent plug. The former provided ade¬ 
quate data on the erosion of gun steel in the course 


a This chapter is based in large measure on a comprehensive 
NDRC report 123 that contains details about the tests of 
erosiveness of propellants. 

b Instructor in Department of Metallurgy, Harvard Uni¬ 
versity. (Present address: Department of Engineering Sciences 
and Applied Physics, Harvard University.) 


of less than 500 rounds; the latter provided such data 
in the course of 15 rounds. Agreement between the 
results of the two methods was in general quite good, 
although some interpretative difficulty was found in 
the case of the vent-plug results. 

Both methods led to orders of erosiveness among 
the various powders tested which were in substantial 
agreement. Generally speaking, the powders were 
found to fall into three groups: those of low erosive¬ 
ness, those of intermediate erosiveness and those of 
high erosiveness. The ordering of the powders within 
the groups tended to differ somewhat according to 
the method of testing. 

It was found that under conditions giving a con¬ 
stant ballistic level, the observed erosion, particularly 
the measured dimensional change in a barrel, is re¬ 
lated to the flame temperature of the powder or to 
the heat input to the barrel during firing. The form 
of the relation is such that a greater and more rapidly 
increasing erosion rate accompanies higher flame 
.temperatures. Thus, under constant known firing 
conditions, powders may be classified on a scale of 
relative erosiveness according to their flame temper¬ 
atures. 

These results should not be taken to mean that the 
high temperature alone, through melting of the bore 
surface, provides the erosion mechanism. Rather 
they mean that temperature is the general regulator 
of whatever erosion mechanisms do operate, such as 
chemical attack by the powder gases and deformation 
and abrasion by the bullets. X-ray diffraction and 
metallographic studies made of the barrels tested in 
this program shed considerable light on these mech¬ 
anisms and their dependence on temperature. They 
supplied much of the data upon which the discussion 
of Chapters 12 and 13 is based. 

It was considered possible early in the program of 
Division 1, as was mentioned in Section 14.1, that 
certain studies might indicate ways in which propel¬ 
lants could be modified in order to mitigate erosion. 
The results of investigations described in this and the 
three preceding chapters indicated that some increase 
in gun life might be possible by only a slight altera¬ 
tion of powder compositions but that no great im¬ 
provement could be expected in the case of steel guns 


308 


CONFIDENTIAL 



METHODS OF TEST 


309 


unless powder compositions were radically changed. 
The modification of propellants is discussed at the end 
of this chapter in Section 15.6. 

15 2 METHODS OF TEST 

15-21 Caliber .50 Erosion Testing Gun 
Gun Assembly 

The caliber .50 erosion-testing gun, described in 
Section 11.2.1, consists essentially of a monobloc, 
45-in. caliber .50 heavy barrel with the origin of rifling 
and bullet seat shaped to receive a 20-mm cartridge 
case necked down to hold a caliber .50 bullet. The 
barrels used in the gun in the testing of different 
propellants were monobloc barrels made of WD 
4150 modified gun steel. Chemical analyses of two 
representative barrels are given in Table 1. 


Table 1 . Chemical analyses of two representative bar¬ 
rels used in the caliber .50 erosion-testing gun in per 
cent. 


Element 

Barrel D-31 

Barrel D-72 

Carbon 

0.56 

0.57 

Manganese 

0.54 

0.54 

Phosphorus 

0.016 

0.016 

Sulfur 

0.019 

0.018 

Silicon 

0.25 

0.25 

Copper 

0.08 

0.08 

Nickel 

0.19 

0.18 

Chromium 

0.96 

0.96 

Vanadium 

0.004 

0.004 

Molybdenum 

0.25 

0.25 

Tin 

0.013 

0.013 

Aluminum 

0.05 

0.05 


Bullets 

Two types of bullets were fired: pre-engraved and 
artillery-type, both of which are described in Section 
11.2.1. The use of these two types of bullets per¬ 
mitted, by comparison, a considerable separation of 
the erosive effects of the hot gases from those of en¬ 
graving stresses and friction. Thus, a good measure 
of the erosion caused by a given propellant alone 
could be deduced. 

Conditions of Test 

Three different series of tests of various powders 
were carried out with the erosion-testing gun. Two of 
the series involved the use, respectively, of pre-en¬ 


graved and artillery-type bullets. The third series 
was carried out at a higher ballistic level with pre¬ 
engraved bullets and included a study of erosion after 
the firing of 1, 5,10, and 70 rounds, respectively, for 
each powder. 

Ballistic Levels. Tests with pre-engraved bullets 
were made with the charges adjusted to give muzzle 
velocities of 3,300 ± 50 fps with each powder. This 
velocity was the highest attainable with a full case of 
the powder having the lowest potential. 

Tests with artillery-type bullets were made with 
the charges adjusted to give maximum powder pres¬ 
sures of 56,000 to 58,000 psi (copper) with each 
powder. 

Tests at higher ballistic level with pre-engraved 
bullets were made with charges adjusted to give 
muzzle velocities of 3,600 fps whenever possible. 

Firing Schedules. The same firing schedule was 
followed in all tests. The rate of fire was 4 rounds per 
minute, the rounds being divided into three groups: 
10 rounds for measurement of pressure and velocity, 
55 rounds for erosion, and 5 rounds for bullet re¬ 
covery. This cycle was repeated until the end of each 
test. 

Measurements 

Pressure and Velocity. As shown in Figure 1 of 
Chapter 11, the chamber of the erosion-testing gun 
is fitted with a copper crusher gauge for the measure¬ 
ment of maximum powder pressure. The accuracy of 
this gauge is estimated at 5 per cent, but it is recog¬ 
nized that in the determination of pressure, the values 
given by such a gauge are considerably lower than 
the true pressure. Increasing the value of the “copper” 
pressure by 20 per cent makes it approximately equal 
to the pressure as measured by a piezoelectric gauge. 

Measurements of velocity were made by means of 
two screens connected to an Aberdeen-type chrono¬ 
graph. The velocities reported are instrumental ve¬ 
locities at 26 ft from the muzzle and it is estimated 
that the deviation of any single velocity from the 
mean is no greater than 0.5 per cent. 

Erosion Rate. The erosion of the gun barrel was 
detected or measured in various ways: by dimen¬ 
sional changes in land and groove diameters, by 
changes in maximum pressure and muzzle velocity, 
by changes in accuracy, and by the changes in the 
engraving observed on recovered bullets. 

The land and groove diameters were measured by 
two series of plug gauges. Those for the lands ranged 


CONFIDENTIAL 







310 


EROSION OF GUN STEEL BY DIFFERENT PROPELLANTS 


from 0.490 in. to 0.516 in. in steps of 0.002 in. while 
those for the grooves were inversely rifled and ranged 
from 0.511 in. to 0.529 in. in steps of 0.002 in. With 
these gauges, land and groove profiles were obtained 
at various stages during the tests. 

Pressure and velocity measurements were made 
during the first 10 rounds of each group of 70 rounds 
of the firing schedule. 

The accuracy was determined at various stages by 
calculating the mean radius of dispersion on targets 
made by the 55 erosion rounds at 100 ft from the 
muzzle. 

Bullet recovery was made in sawdust and band 
diameter and width of engraving were measured. Any 
body engraving was also noted. 

Limited amounts of the special powders tested pre¬ 
cluded any exhaustive study of the reproducibility of 
erosion as measured by these procedures. However, 
several pairs of check tests were made and they gave 
results in excellent agreement with one another. 

Heat Input. The heat input to the bore was cal¬ 
culated from temperature-time curves which were 
recorded with thermocouples and Speedomax record¬ 
ers. Special barrels with wall sections machined down 
to ^ in. near the origins of rifling were used in making 
these measurements and the thermocouples were 
placed at several positions along the thin wall, as 
described in Section 5.3.4. Heat input values were 
measured 71 for the firing of each of several rounds 
with loading conditions of the various powders the 
same as in erosion tests. 

Metallographic Changes. Sections of the fired bar¬ 
rels were examined metallographically in order to 
detect changes in the condition of the bore surface 
and in the microstructure of the gun steel. These 
studies were more related to the mechanism of the 
erosion than to the amount of erosion. 

X-Ray Studies. The bore surfaces of the fired bar¬ 
rels were studied by x-ray diffraction in order to iden¬ 
tify the erosion products remaining on the bore sur¬ 
faces. These studies were also related to the erosion 
mechanism rather than to the amount of erosion. 


15,2,2 Erosion Vent Plug 

An independent experimental study of the erosive¬ 
ness of many of the powders which were tested in the 
erosion-testing gun was made using the vent plug 
technique described in Section 11.2.3. Results ob¬ 
tained by these two dissimilar methods of study pro¬ 


vide corroborative checks on the behavior of the 
propellants. 

Apparatus 

The explosion vessel, which could be electrically 
fired, contained a spherical chamber which commun¬ 
icated with both a detachable erosion vent of cylindrical 
shape and a piezoelectric pressure gauge. The vent 
was closed at its muzzle end with a calibrated brass 
rupture disk. The explosion vessel had a chamber 
volume of 24.3 cc and, together with its fittings, was 
made of NE 9450 steel. 

Vent Plugs. The cylindrical plugs for the powder 
testing program were made of SAE 4140 gun steel, 
heat treated to a hardness of Rockwell C-18 as meas¬ 
ured at the extremities of the cylinder axis before 
drilling of the vent. The length of the plugs was % in. 
Since preliminary work showed that the erosion of 
vents increased with the roughness of the interior 
vent surface, it was decided to make that surface the 
smoothest practicable. This aim was accomplished 
by first reaming the vents 0.0005 in. undersize and 
then lapping the surface to final dimensions. 

Rupture Disks. The rupture disks were stamped 
from a single sheet of yellow brass having a thickness 
of 0.042 ± 0.0017 in. and very uniform hardness. The 
rupture area was 0.049 sq in. These disks were cali¬ 
brated quite simply by firing increasing charges of 
FNH-M2 powder and observing the pressure-time 
curves. It was found that rupture of the 0.042-in. 
disk occurred at 25,000 ± 1000 psi and the determi¬ 
nation was later confirmed through the use of an 
automatic recording of the instant of rupture on the 
pressure-time curves. The device consisted of a fine 
wire stretched before the rupture disk which broke at 
the instant of rupture and in so doing caused a sudden 
change in the potential across the recording oscillo¬ 
graph. 

Procedure 

Ideal operation of the vent-plug apparatus would 
permit each shot with any powder to build up powder 
gases to the same maximum pressure, whereupon 
these gases would suddenly sweep through the vent. 
In this manner, erosion in the vent would be quite 
independent of the burning characteristics of the 
various powders and would reflect only the effect of 
the gases themselves. In practice this ideal is difficult 
to achieve primarily because of small variations in 


CONFIDENTIAL 



METHODS OF TEST 


311 


both the maximum pressures and rupture pressures 
from round to round. Thus, if the charge is so ad¬ 
justed that the maximum pressure is about the same 
as the rupture pressure, a large number of failures of 
disk rupture are encountered. This difficulty was 
avoided by using charges of the different powders 
which yielded maximum pressures about 3,000 psi 
higher than the rupture pressure. Only a very small 
overlap of burning and erosion times was thus per¬ 
mitted. 

Charge Determination. Determination of the proper 
charges of the various powders tested was made by 
trial and error. A tendency for the peak pressure to 
decrease with increasing vent size was noted, but it 
was felt that the order of magnitude of the variation 
was too small to have an appreciable effect on the 
erosion. 

Vent Sizes. Ordinarily 4 degrees of severity of test 
were provided by the use of vents of four different 
internal diameters: %, 34 % and 34 in., in order of 
increasing severity of test. RDX powders were tested 
with vents of three different sizes: 34 and 34 in. 

Tolerances were held to 0.0005 in. for the larger vents 
and 0.001 in. for the 34 _ in. vent. 

Measurements. Erosion was produced by the firing 
of repeated rounds with various powders at each of 
the levels of severity afforded by the different vent 
sizes. Two identical vent plugs of each size were tested 
with each powder, the plugs being alternated so as to 
subject them as nearly as possible to equal treatment. 
These two tests provided checks on one another. 
Each plug was weighed to the nearest milligram at 
the completion of 1, 5, 10 and 15 rounds after having 
been gently scraped, swabbed with acetone and with 
alcohol, and carefully dried. Erosion was reported in 
terms of the weight loss of each plug. 

Validity of Measurements. It was found that the 
deviation from the mean weight loss of the two check 
runs in each pair of tests was quite small, being of the 
order of 10 per cent for the vent tests and 

possibly even smaller for the smaller vents. These 
deviations were also smaller after 15 rounds than 
after 5 or 10 rounds. It was inferred, therefore, that 
reproducibility of these tests was adequate. 

The temperature of the explosion chamber varied 
somewhat from test to test, but it was determined by 
direct experiment that these variations had negligible 
effect on the results. 

Attention was paid to the possibility that the pow¬ 
der gases might be cooled considerably before coming 
in contact with the vent plugs and that the extent of 


such cooling might differ widely in different tests. 0 
Accordingly the rate of pressure change with time 
just after the end of burning of a slow-burning pow¬ 
der was measured at various densities of loading. 
From these results it was found that the cooling rate 
for the powder gas was such that its temperature 
could not have fallen more than 10 C during the 
4-msec period required for the burning of this powder 
prior to the rupture of the disk in one of the erosion 
vent experiments. Therefore it was assumed that gas 
temperatures for all the powders at the beginning of 
erosion were essentially equal to the adiabatic flame 
temperature. 

At any stage after rupture, both experimental and 
theoretical results indicate that all gas temperatures 
are similarly related to the respective adiabatic flame 
temperatures of their powders. It follows, then, that 
direct comparison of weight losses of the vent plugs 
permits an accurate determination of the relative 
erosiveness of the various powders which is depend¬ 
ent only on the flame temperatures and chemical 
properties of those powders. 

15,2,3 Powders Tested 

A total of 21 different powders was tested in either 
the caliber .50 erosion-testing gun or the erosion vent 
plug, most powders being tested in both. Character¬ 
istics of the powders and the methods of testing them 
are given in Table 2 where they are listed in order of 
increasing adiabatic flame temperature. The powders 
were all of standard lots except for those for which 
the experimental lot numbers are given in column 2. 
The latter were supplied by Division 8, NDRC. The 
adiabatic flame temperatures given in column 3 were 
calculated according to the method described in Sec¬ 
tion 2.4.3. The numbers in column 4 designate the 
types of test to which the powders were subjected, as 
follows: 

1. Erosion-testing gun with pre-engraved bullets. 

2. Erosion-testing gun with artillery-type bullets. 

3. Erosion-testing gun with pre-engraved bullets at 
high ballistic level. 

4. Erosion vent plug. 

c In experiments conducted at Explosives Research Labor¬ 
atory under the supervision of Division 8, NDRC, some indi¬ 
cation was obtained “that cooling during burning was the 
same for a large granulation as for a small granulation of the 
same powder, because the integrated turbulence was the same 
in each case.” (Personal communication from Dr. J. F. Kin¬ 
caid, formerly of the Explosives Research Laboratory, after 
having reviewed the manuscript of this chapter, July 24, 1946.) 


CONFIDENTIAL 




312 


EROSION OF GUN STEEL BY DIFFERENT PROPELLANTS 


Table 2. Characteristics and methods of test of the various powders.* 


Powders 

Lot No. 

Flame 

temperature 

(K) 

Tests 

Specific heat at 
constant volume 
(cal/gram) 

Impetus of 
powder f 
(ft-lb/lb) 

RDX 

5060 

2320 

1 



RDX 

6079 

2465 

4 



Cordite N 


2469 

1,2 

0.3616 

335,000 

FNH/P 


2480 

1,2,4 

0.3456 

310,000 

FNH-M1 


2483 

1, 2, 3, 4 

0.3456 

310,000 

RDX 

5061 

2560 

1 



RDX 

6080 

2611 

4 



NH-M1 

(3727) t 

2651 

4 



NH-M1 

4974 

2696 

4 



RDX 

6081 

2713 

4 



NH-M1 


2807 

1,2 

0.3425 

328,010 

Pyro 


2814 

1,2,4 

0.3692 

327,490 

Cordite NQ 


2864 

1,2 

0.3588 

363,000 

IMR 


2938 

1, 2, 3, 4 

0.3381 

344,000 

RDX 

6082 

3003 

4 



RDX 

5059 

3080 

2 



FNH-M5 


3268 

1,2 

0.3445 

362,750 

RDX 

6083 

3302 

4 



FNH-M2 


3563 

1, 2, 3, 4 

0.3439 

384,000 

Ballistite (40% NG) 


3945 

3 

0.3393 

410,400 

Ballistite (60% NG) 


4300 

3 

0.3390 

441,400 


* This table has been compiled from Tables I, II, XXIX, and XXXVa of the NDRC report 123 on this subject, which also gives detailed descriptions 
of the powders. 

t Defined in Section 3.2.3. 

J Powder for 37-mm gun M1916; standard lot No. 3727. 


Table 3. Firing conditions for different powders in the erosion-testing gun. 


Powder 

Web 

(in.) 

Charge 

(grains) 

Loading 
density 
(g per cu cm) 

Maximum 
pressure 
[psi (copper)] 

Initial 

velocity 

(fps) 

Rounds 

fired 


A. Charges to give initial muzzle velocities of about 3,300 fps with pre-engraved bullets. 


RDX-5060 

0.019 

385 

0.762 

47,800 

3,367 

507 

Cordite N 

0.0186 

415 

0.822 

54,835 

3,302 

480 

FNH-M1 

0.0200 

420 

0.832 

55,525 

3,335 

500 

FNH/P 

0.0234 

425 

0.842 

52,123 

3,300 

332 

NH-M1 

0.0262 

415 

0.822 

55,600 

3,365 

305 

Cordite NQ 

0.0181 

345 

0.683 

57,509 

3,353 

510 

RDX-5061 

0.025 

380 

0.752 

49,710 

3,356 

500 

Pyro 

0.0245 

405 

0.802 

56,100 

3,288 

360 

IMR 

0.0318 

425 

0.842 

46,880 

3,326 

500 

FNH-M5 

0.0338 

360 

0.713 

46,460 

3,305 

307 

FNH-M2 

0.0350 

350 

0.693 

46,600 

3,324 

220 


B. Charges to give 

maximum 

pressures of about 57,000 psi (Cu) with artillery-type bullets. 


Cordite N 

0.0255 

440 

0.893 

57,920 

3,290 

557 

FNH-M1 

0.0241 

425 

0.862 

56,375 

3,252 

500 

FNH/P 

0.0240 

425 

0.862 

57,615 

3,286 

355 

NH-M1 

0.0294 

425 

0.862 

56,120 

3,280 

360 

Pyro 

0.0285 

425 

0.862 

53,000 

3,210 

360 

Cordite NQ 

0.0244 

420 

0.853 

55,685 

3,434 

514 

IMR 

0.0338 

425 

0.862 

56,700 

3,428 

515 

FNH-M5 

0.0454 

385 

0.782 

56,920 

3,437 

377 

FNH-M2 

0.0461 

400 

0.812 

57,935 

3,458 

220 


C. Charges to give initial velocities of 3,600 fps with pre-engraved bullets. 


IMR 

0.0318 

476 

0.941 

58,000 

3,600 


FNH-M2 

0.0350 

395 

0.782 

58,000 

3,600 


40% NH 

0.0490 

390 

0.762 

58,000 

3,600 


60% NG 

0.0395 

330 

0.654 

58,000 

3,600 



CONFIDENTIAL 































TESTS IN CALIBER .50 EROSION-TESTING GUN 


313 


Column 5 gives the nominal specific heat at constant 
volume and column 6 the nominal impetus of the 
powder. Both were calculated according to the 
method of Hirschfelder. 26 

The loading conditions in the various tests varied 
widely so as to maintain the desired ballistic level 
with each of the different powders. Firing data are 
given in the three parts of Table 3 for the three series 
of tests in the erosion-testing gun and in Table 4 for 
the tests with the erosion vent plug. 


Table 4. Firing conditions for different powders in the 
erosion vent plug chamber. Charges to give maximum 
pressures of 28,000 psi (piezo). 


Maximum pressure 
Web Charge (for vent) 

Powder (in.) (g) [psi (piezo)] 


RDX-6079 . 4.0 . 

FNH/P 0.0106 4.4 29,000 

FNH-M1 0.0105 4.4 28,500 

RDX-6080 . 3.9 . 

NH-M1* 0.0120 4.18 28,500 

NH-M1 0.011 4.2 29,000 

RDX-6081 . 3.8 . 

Pyro 0.0089 4.0 29,000 

IMR 0.0099 3.97 28,000 

RDX-6082 . 3.7 . 

RDX-6083 . 3.4 . 

FNH-M2 0.015 3.6 28,000 


* Powder for 37-mm gun M1916; standard lot No. 3727. 


i5.3 TESTS IN CALIBER .50 

EROSION-TESTING GUN 

15 31 Dimensional Erosion Rate 

At frequent intervals in the course of the testing of 
each powder in the erosion-testing gun the barrel was 
removed from the gun assembly and measured in¬ 
ternally with plug gauges. These measurements were 
expressed in terms of the observed increase in land 
diameter A L and in groove diameter AG at various 
points along the barrel. Plotting both A L and AG 
versus distance from the origin of rifling gave a 
double erosion-distribution curve for each stage of 
the test. 

The progress of erosion at each point along the 
barrel was easily obtained by comparing the erosion- 
distribution curves at different stages during the test. 
Data taken from these curves permitted the plotting 
of A L and AG versus number of rounds for any given 
position along the barrel. 


Distribution of Erosion 

Pre-engraved Bullet Series. Separate distributions 
of land and groove erosion as measured by plug 
gauges after completion of the various tests are plot¬ 
ted in Figures 1 and 2, respectively. The tests were 
all carried out at initial muzzle velocities of 3,300 fps, 
and the numbers of rounds fired in each test are given 
in the legends. 

It will be noted that no erosion was detected in the 
barrels fired with very cool powders and that the 
amount of erosion increased generally with the flame 
temperature of the powder. In every case, erosion 
was most severe near the origin of rifling and fell off 
rapidly toward the muzzle. The extension of meas¬ 
ured erosion muzzleward increased with the flame 
temperature of the powder. 

Artillery-Type Bullet Series. Similar land and groove 
erosion distribution curves for all tests with artillery- 
type bullets are plotted in Figures 3 and 4, respec¬ 
tively. These tests were carried out at maximum 
pressures of about 57,000 psi (copper). The numbers 
of rounds fired are given in the legends. 

It will be noted that in these tests erosion was 
detected even when very cool powders were fired. As 
before, erosion increased in amount and advanced 
muzzleward as the flame temperature of the powder 
was increased. In all cases the greatest erosion was 
found near the origin of rifling. 

High Ballistic Level Series. Little attention was 
paid to distribution of erosion in this series, for which 
the initial muzzle velocity was 3,600 fps. Pre-engraved 
bullets were fired. Star gauge measurements were 
made only at 34 in- and 1 in. beyond the origin of 
rifling to determine the progress of erosion after the 
firing of 1, 5, 10, and 70 rounds. 


Progress of Erosion 

Pre-engraved Bullet Series. The progress of land 
erosion was obtained from the distribution curves by 
observing the measured land diameter increase at 
three positions along the barrel at various stages of 
each test. These data are listed in Table 5. The prog¬ 
ress of erosion 34 in. from the origin of rifling is 
plotted for the various tests in Figure 5. 

It will be seen that the erosion rate is a constant in 
these tests; that is, the increase in land diameter A L 
is a linear function of the number of rounds fired. The 
barrels fired with cooler powders did not start to 
erode at once, however, so that their curves intersect 


CONFIDENTIAL 


















314 


EROSION OF GUN STEEL BY DIFFERENT PROPELLANTS 


0.518 



0.490 


5 6 7 8 9 10 II 12 

DISTANCE FROM ORIGIN OF RIFLING IN INCHES 


Figure 1. Distribution of land erosion for various powders fired with pre-engraved bullets. (This figure has appeared 
as Figure 8 in NDRC Report No. A-451.) 


0.527 


0.525 


| 0.523 


5 0.521 

f— 

UJ 

5 

< 0.519 

Q 

UJ 

O 

3 0.517 
o 

UJ 

o 0.515 
o 

K 

«> 

0.513 


0.511 

0 1 2 3 4 5 6 7 8 9 10 II 12 13 14 15 

DISTANCE FROM ORIGIN OF RIFLING IN INCHES 





"RDX-5060 

AFTER 

507 

ROUNDS 



(lb 

FNH-M1 

AFTER 

500 

ROUNDS 



CORDITE N 

AFTER 

480 

ROUNDS 




_FNH/P 

AFTER 

332 

ROUNDS 


o 

(2) 

NH 

AFTER 

305 

ROUNDS 

— 

X 

(3) 

CORDITE NQ 

AFTER 

510 

ROUNDS 


V 

(4) 

RDX-5061 

AFTER 

500 

ROUNDS 


A 

(5) 

PYR0 

AFTER 

360 

ROUNDS 

_ 

□ 

(6) 

IMR 

AFTER 

500 

ROUNDS 


® (7) 

FNH-M5 

AFTER 

307 

ROUNDS 


B 

(8) 

FNH-M2 

AFTER 

220 

ROUNDS 

_ 


Figure 2. Distribution of groove erosion for various powders fired with pre-engraved bullets. (This figure has appeared 
as Figure 9 in NDRC Report No. A-451.) 


Table 5. Pre-engraved bullet series. Increase in land diameter found at locations lying y 2 , 1, and 2 in. beyond the origin 
of rifling after the firing of 140, 280, and 350 rounds. 


Powder 

After 140 rounds 

After 280 rounds 

After 350 rounds 

34 in. 

(in. X 10~ 3 ) 

1 in. 

(in. X 10~ 3 ) 

2 in. 

(in.XlO- 3 ) 

34 in. 
(in. X 10~ 3 ) 

1 in. 

(in.XlO -3 ) 

2 in. 

(in.XlO- 3 ) 

34 in. 

(in.XlO- 3 ) 

1 in. 

(in.XlO- 3 ) 

2 in. 

(in.XlO- 3 ) 

RDX-5066 

0 

0 

0 

0 

0 

0 

0 

0 

0 

Cordite N 

0 

0 

0 

0 

0 

0 

0 

0 

0 

FNH-M1 

0 

0 

0 

0 

0 

0 

0 

0 

0 

FNH/P 

0 

0 

0 

0 

0 

0 

0 

0 

0 

NH-M1 

0.5 

0 

0 

2.6 

1.1 

0 

3.7 

1.7 

0 

Cordite NQ 

1.8 

0.8 

0 

3.9 

1.7 

0.2 

5.0 

2.2 

0.8 

RDX-5061 

1.5 

0.7 

0.6 

3.9 

2.2 

1.4 

5.0 

2.9 

1.5 

Pyro 

2.5 

0.6 

0 

6.6 

2.5 

0 




IMR 

3.0 

1.3 

0 . 

6.8 

4.8 

2.5 

8.9 

6.7 

3.8 

FNH-M5 

14.4 

12.3 

7.6 

32.6 

22.8 

16.6 




FNH-M2 

17.4 

13.2 

7.5 

.... 


.... 





CONFIDENTIAL 


























TESTS IN CALIBER .50 EROSION-TESTING GUN 


315 


0.518 


0.514 

V) 

UJ 

g 0.510 


“ 0.506 

UJ 

t— 

UJ 

2 0.502 
< 

Q 

UJ 

o 0.498 

< 

o 

§ 0.494 
< 

_i 

0.490 


0.486 

0 1 2 3 4 5 6 7 8 9 10 II 12 13 14 15 

DISTANCE FROM ORIGIN OF RIFLING IN INCHES 


I I I I I I 



CORDITE 

FNH-M1 

FNH/P 

NH 

PYRO 

CORDITE 

IMR 

FNH-M5 

FNH-M2 


N AFTER 
AFTER 
AFTER 
AFTER 
AFTER 
NQ AFTER 
AFTER 
AFTER 
AFTER 


557 ROUNDS 
504 ROUNDS 
355 ROUNDS 
360 ROUNDS 
360 ROUNDS 

514 ROUNDS 

515 ROUNDS 
377 ROUNDS 
220 ROUNDS 


1 I 1 I I 


Figure 3. Distribution of land erosion for various powders fired with artillery-type bullets. (This figure has appeared 
as Figure 13 in NDRC Report No. A-451.) 



I I I I I I 


CORDITE 

N AFTER 

557 

ROUNDS 


FNH-M1 

AFTER 

504 

ROUNDS 

— 

FNH/P 

AFTER 

355 

ROUNDS 


NH 

AFTER 

360 

ROUNDS 


PYRO 

AFTER 

360 

ROUNDS 

— 

CORDITE NQ AFTER 

514 

ROUNDS 


IMR 

AFTER 

515 

ROUNDS 


FNH-M5 

AFTER 

377 

ROUNDS 


FNH-M2 

AFTER 

220 

ROUNDS 



__i_ >—i i ' 

4 5 6 7 8 9 10 

DISTANCE FROM ORIGIN OF RIFLING IN INCHES 


12 


13 


14 


15 


Figure 4. Distribution of groove erosion for various powders fired with artillery-type bullets. (This figure has appeared 
as Figure 14 in NDRC Report No. A-451.) 


Table 6. Artillery-type bullet series. Increase in land diameter found at locations lying l /%, 1, and 2 inches beyond the 
origin of rifling after the firing of 140, 280, and 350 rounds. 


Powder 

After 140 rounds 

After 280 rounds 

After 350 rounds 

Yi in. 1 in. 2 in. in. 1 in. 2 in. 3^ in. 1 in. 

(in. X 10 -3 ) (in. X 10~ 3 ) (in. X 10~ 3 ) (in. X 10 -3 ) (in. X 10 -3 ) (in. X 10 -3 ) (in. X 10 -3 ) (in. X 10 -3 ) 

2 in. 

(in. X 10- 3 ) 

Cordite N 

0.2 

0 

0 

2.3 

0 

0 

2.8 

0 

0 

FNH-M1 

1.4 

0 

0 

4.5 

3.6 

2.3 

5.4 

4.6 

3.2 

FNH/P 

4.2 

3.1 

1.1 

6.1 

4.2 

2.9 

6.6 

4.2 

3.2 

NH-M1 

4.4 

2.2 

0 

6.3 

3.3 

0 

6.6 

3.0 

0 

Pyro 

5.2 

2.1 

0 

8.5 

3.0 

0 

9.5 

3.5 

0 

Cordite NQ 

6.1 

4.2 

2.1 

11.4 

1.8 

3.7 

14.0 

9.5 

4.4 

IMR 

9.5 

7.7 

4.6 

14.7 

12.3 

7.5 

16.2 

13.7 

8.6 

FNH-M5 

14.8 

12.9 

7.8 

33.0 

24.9 

15.8 

40.0 

31.8 

19.6 

FNH-M2 

30.0 

24.8 

17.5 




.... 




CONFIDENTIAL 






















316 


EROSION OF GUN STEEL BY DIFFERENT PROPELLANTS 



ROUNDS FIRED N 


Figure 5. Progress of erosion at ^ in. beyond the origin of rifling for various powders fired with pre-engraved bullets. 
(This figure has appeared as Figure 10 in NDRC Report No. A-451.) 



Figure 6. Progress of erosion at in. beyond the origin of rifling for various powders fired with artillery-type bullets. 
(This figure has appeared as Figure 15 in NDRC Report No. A-451.) 


CONFIDENTIAL 

















TESTS IN CALIBER .50 EROSION-TESTING GUN 


317 


the abscissa. The erosion rate increases as the flame 
temperature of the powder increases. 

Artillery-Type Bullet Series. The progress of land 
erosion was obtained exactly as before, the data being 
listed in Table 6 and being plotted for points located 
Y 2 in. from the origin of rifling in Figure 6. 

The erosion rate in these tests will be seen to be no 
longer constant, but rather to decrease as the test was 
prolonged. As before, erosion did not start at once in 
the barrels fired with cooler powders so that their 
curves intersect the abscissa. The erosion rate at any 
stage again appears to increase as the flame tempera¬ 
ture of the powder increases. 

High Ballistic Level Series. Progress of land erosion 
data at a comparable position along the barrel (Y in* 
from the origin of rifling) was obtained by interpola¬ 
tion of the star gauge data taken at various stages at 
positions lying Y in. an d 1 in. beyond the origin of 
rifling. The results are listed in Table 7 and are plot¬ 
ted in Figure 7. 


Table 7. High ballistic level series. Increase in land 
diameter at Yi inch beyond origin of rifling (interpo¬ 
lated) after the firing of 1, 5, 10, and 70 rounds. 


Powder 

1 Round 
(in. X 10- 3 ) 

5 Rounds 
(in. X 10- 3 ) 

10 Rounds 70 Rounds 
(in. X IO- 3 ) (in. X 10~ 3 ) 

IMR 

0 

1.2 

0.5 

6.0 

FNH-M2 

0 

1.2 

1.5 

21.5 

40% NG 

0.9 

4.5 

7.0 

48.2 

60% NG 

0.6 

4.8 

8.7 

66.5 


As before in the pre-engraved bullet series (at 
moderate ballistic level), the erosion rate appears to 
be constant, but in this series erosion appears to start 
with the first round. The erosion rate increases with 
the flame temperature of the powder. 

Standardized Measure of Erosion Rate 

The erosion rate characteristic of each of the tested 
powders should be reduced to a single index if com¬ 
parisons are to be made easily. This result is easily 
achieved in the case of tests where pre-engraved bul¬ 
lets are fired, for the erosion rate is a constant in 
these tests for any definite distance from the origin of 
rifling. It remains only to note that the erosion rate 
E (in. per round) is given by equation (1), 


where A L is increase in land diameter in inches at 


some definite distance from the origin of rifling; N is 
total number of rounds fired; and n is number of 
rounds fired before measurable erosion occurs. E is, 
of course, the slope of the progress of erosion curve at 
the location in question and n is the intercept on the 
abscissa. 

The situation is considerably complicated in the 
case of the firing of artillery-type bullets, for the 
erosion rate at any given location is not constant. 
The reason it is not constant is related to the role 
played in erosion by the considerable engraving- 
stresses which accompany the firing of these bullets. 



Figure 7. Progress of erosion at 3^ in. beyond the 
origin of rifling for propellants of high ballistic level. 
(This figure has appeared as Figure 19 in NDRC Re¬ 
port No. A-451.) 

In a new barrel there is large diametrical interference 
between the lands and the rotating bands which leads 
to very severe land wear. This wear, in turn, reduces 
the land height so that the interference becomes less. 
Thus, the rate of this mechanical type of land wear is 
initially large, but decreases as wear proceeds. Super¬ 
posed upon it is the erosion caused by the powder 
gases, the sum being the measured erosion. 

The erosion rate obtained with artillery-type bul¬ 
lets is properly expressed as the slope of the erosion- 
progress curve drawn for some given location in the 
barrel at a given stage of the firing. An average value 
of this rate might also be reported, but such a figure 
would be without mechanistic significance. 

The best location at which the erosion rate may be 
specified is where the erosion is most severe; that is, 
near the origin of rifling. For this reason, the location 
Y 2 in. muzzleward from the origin of rifling has been 
selected. The curves showing progress of erosion for 
this location have been plotted in Figures 5, 6, and 7. 

The best stage of the test at which to specify the 
erosion rate is probably toward the end of the test, 


CONFIDENTIAL 











318 


EROSION OF GUN STEEL BY DIFFERENT PROPELLANTS 


for at that stage the relative effect of engraving stres¬ 
ses is less and the results are more representative of 
the effect of powder gases alone. Alternatively, the 
erosion rate might be specified at such stages that the 
observed reduction in land height was the same in 
each of the various tests. It is not certain, however, 
that the proportion of total erosion resulting from 
engraving stresses is always the same at given land 
height independent of other factors. 

The erosion rates for the various powders d are tab¬ 
ulated in Table 8 by series. The listings for the pre- 


Table 8. Erosion rates observed with various powders 
in the three firing series. 


Powder 

Flame 

temper¬ 

ature 

(K) 

Pre- Artillery- 

engraved type 

bullet series bullet series 
(in. per rd (in. per rd 

x 10- 5 ) x io- 6 ) 

High 
ballistic 
level series 
(in. per rd 

X 10“ 5 ) 

RDX-5060 

2320 

0 



Cordite N 

2469 

0 

0.7 


FNH/P 

2480 

0 

0.8 


FNH-M1 

2483 

0 

1.2 


RDX-5061 

2560 

1.66 



NH-M1 

2807 

1.42 

1.0 


Pyro 

2814 

2.60 

2.0 


Cordite NQ 

2864 

1.52 

3.7 


IMR 

2938 

2.73 

4.0 

8.5 

FNH-M5 

3268 

10.22 

8.8 


FNH-M2 

3564 

12.52 

27.0 

30.7 

40% NG 

3945 



68.8 

60% NG 

4300 



95.0 


engraved and the high ballistic level series are, of 
course, values of E found from the equation given 
above in this section. The listings for the artillery- 
type bullet series are slopes of the progress-of-erosion 
curves for 3^ in. beyond the origin of rifling at 300 
rounds. Values for the two hottest powders were 
found by extrapolating their essentially straight, 
progress-of-erosion curves to 300 rounds. 

It will be noted that there are some minor differ¬ 
ences in order between the erosion rates of the pre¬ 
engraved series and the artillery-type bullet series 
erosion rates. In general, the artillery-type bullet 
values are more in line with the flame temperatures 
of the powders and their greater magnitude has per¬ 
mitted ordering of the powders of low erosiveness 
which produced no measurable erosion with pre¬ 
engraved bullets. The relation between the erosion 

fl At Picatinny Arsenal the quickness of standard Pyro and 
FNH powders has been determined, 316 and thermochemical 
and physical tests 317 of nitroguanidine powders have been per¬ 
formed. 


rate and the adiabatic flame temperature is discussed 
in Section 15.5.2, where the data of Table 8 are 
presented graphically in Figure 9. 

In any event, the groupings of powders are more 
important than their order within a group. The aver¬ 
age of a much larger number of tests would be re¬ 
quired to classify the powders in an absolute order of 
increasing erosiveness. The following groupings are 
evident, except that it is uncertain whether NH-M1 
should be in the first or second group: 

RDX-5060, Cordite N, FNH/P, FNH-M1, 

NH-M1 (?) < RDX-5061, Pyro, Cordite NQ, 

< IMR < FNH-M5 < FNH-M2 < 40% NG 

< 60% NG. 

15,3,2 Heat Input to the Barrel 

The heat input to the barrel was measured accord¬ 
ing to the method described in Section 5.3.4 at two 
positions along the barrel: 1.25 in. and 4.50 in. beyond 
the origin of rifling. These determinations were not 
made during the tests used to find dimensional ero¬ 
sion, but on specially machined barrels which were 
fired with equivalent powder charges. Both pre-en- 
graved and artillery-type bullets were fired. 

While it is known that the generation of heat 
accompanying the engraving of artillery-type bullets 
results in a preheating of the bore, the magnitude of 
the preheating was too small to be detected by these 
means. Accordingly, the heat input was found to be 
characteristic of the powder, but independent of the 
bullets fired. In Table 9 the measured heat input 
accompanying the firing of the various powders is 


Table 9. Flame temperatures of the various powders 
and heat input as measured at 1.25 in. beyond the origin 
of rifling of caliber .50 erosion-testing gun. 


Powder 

Flame 

temperature 

(K) 

Heat 

input 

(cal per sq cm) 

RDX-5060 

2320 

11.7 

Cordite N 

2469 

11.8 

FNH/P 

2480 

11.8 

FNH-M1 

2483 

12.2 

RDX-5061 

2560 

12.8 

NH-M1 

2807 

13.3 

Pyro 

2814 

13.8 

Cordite NQ 

2864 

12.7 

IMR 

2938 

14.4 

RDX-5059 

3080 

14.4 

RNH-M5 

3268 

16.4 

FNH-M2 

3563 

17.1 

40% NG 

3945 

18.5 

60% NG 

4300 

19.2 


CONFIDENTIAL 
















TESTS IN CALIBER .50 EROSION-TESTING GUN 


319 


listed along with the flame temperatures of the 
powders. 

With the exception of the anomalous value for 
cordite NQ powder, the heat inputs appear to agree 
with the flame temperatures of the powders. Indeed, 
as is shown in Section 15.5.3, the relation between 
flame temperature and heat input appears to be lin¬ 
ear except for the hotter powders. These latter are 
known to cause considerable melting of the bore sur¬ 
face and it seems probable that the absorption of 
latent heat of fusion accounts for the somewhat low 
heat input observed during their firing. 

The correlation between heat input and erosion 
rate is discussed in Section 15.5.2. 

15 3 3 Metallographic Changes 

Accompanying Firing 

The gun steel barrels of the powder testing pro¬ 
gram were examined metallographically upon com¬ 
pletion of the firing tests and the details of many 
changes in the nature of the bore surface caused by 
firing, such as those described in Chapters 10 and 12, 
were thus revealed. Outstanding among these changes 
were the cracking of the surface, pebbling of the 
surface and the development of a layer of trans¬ 
formed steel next to the surface. Relationship be¬ 
tween these changes and the flame temperature of 
the powder was established in several aspects. 

Cracking of the Surface 

Surface cracking always developed as a network 
characteristic of the type that has been attributed to 
thermally induced stresses (Section 13.5.3). The mesh 
size generally was less as the powder flame tempera¬ 
ture was lowered. The apparent width of the cracks 
appeared to increase directly with the flame temper¬ 
ature and the number of rounds fired. Their depth 
was rather anomalous, however, being greater in the 
case of cooler powders. 

It seems probable that the incipient melting which 
occurs during the firing of all but the coolest powders 
removes stress raisers in the surface with the result 
that cracks of full depth do not develop. Erosion from 
the surface also reduces the apparent depth of 
cracking. 

Pebbling of the Surface 

The bore surfaces of barrels fired with any but the 


coolest powders had a pebbled appearance caused by 
the rounding of the edges of the cracks by liquefac¬ 
tion, as described in Section 10.5.2. Since the pebbling 
was related to the crack system, it was coarser grained 
the hotter the powder. It has been shown 124 that the 
extent of the pebbling can be correlated with the 
excess of a calculated maximum bore-surface temper¬ 
ature over the melting point of the gun steel, the 
maximum surface temperature being found, as de¬ 
scribed in Section 5.4.1, on the assumption that no 
melting occurs. 

When the hottest powders are fired the very coarse 
pebbling merges smoothly with a general type of 
melting which produces a rippled surface. This type 
of melting is presumably a true melting of gun steel, 
although it is recognized that chemical changes may 
play a role in lowering the melting point of the steel. 

Thermal Transformation of the Surface 

The thermal transformation of a gun steel layer 
next to the bore surface has been discussed in some 
detail in Sections 12.1.2 and 13.2.3. Much of the 
quantitative data which has pointed to a possible 
mechanism of the formation of this layer 124 was ob¬ 
tained from a metallographic study of the barrels 
fired in the powder testing program. Particularly, the 
rate of formation and the relation between dimen¬ 
sions of the transformed layer and the flame temper¬ 
ature of the powder were established through this 
study. 

Distribution of the Transformed Layer. The total 
altered layer on a bore surface is properly divided 
into at least two distinct parts: a chemically altered, 
outer layer, to which the term “white layer” is restrict¬ 
ed for the reasons given at the end of Section 12.1.1, 
and an inner layer which has essentially the chemical 
composition of gun steel but which has been ther¬ 
mally altered. The white layer usually has negligible 
thickness compared with the thermally altered layer. 
In the following discussion the thickness reported for 
the transformed layer is actually the total thickness 
of both layers, but it is to be understood that this 
thickness, in the case of the barrels studied in con¬ 
nection with the powder-testing program, is not sig¬ 
nificantly different from that of the thermally altered 
layer alone. 

By virtue of its well-established independence of 
chemical effects, it is to be expected that the thick¬ 
ness of the thermally transformed layer will depend 
on the severity of thermal conditions along the barrel 


CONFIDENTIAL 





320 


EROSION OF GUN STEEL BY DIFFERENT PROPELLANTS 


in which it forms. The expectation is fully confirmed 
by the distribution of the layer in the barrels studied. 
In every case the layer was found to be thickest in 
the neighborhood of the origin of rifling where ther¬ 
mal conditions are most severe and to decrease 
muzzleward. 

This variation presents a minor problem when two 
or more different test barrels are to be compared. The 
best point of reference is at the origin of rifling where 
the layer is thickest, but often the thickness changes 
so rapidly in this region that minor errors in section¬ 
ing the barrel cause large errors in thickness measure¬ 
ment. The problem was solved by determining the 
thickness at different points away from the origin of 



Figure 8. Distribution of transformed layer for dif¬ 
ferent numbers of rounds of double-base (40% NG) 
powder. (This figure has appeared as Plot VI in NDRC 
Report No. A-452.) 

rifling, usually 34 1, 2, and 4 in., drawing a smooth 
curve through these points, and extrapolating to the 
origin itself. Curves of this type are shown in Figure 
8 . 

It was found that the geometry of the projecting 
lands enhanced the heat transfer so that transformed 
layers in the lands were thicker and less sensitive to 
the type of powder than layers in the grooves. The 
layer thickness most characteristic of the firing of a 
given powder was therefore taken as the extrapolated 
thickness in the grooves. 

Layer thicknesses determined in this way were 
found to be independent of the type of bullets or the 
number of rounds fired. They are tabulated in Table 
10 . 

Rate of Formation of the Transformed Layer. The 
barrels fired in the series of tests at high ballistic level 


Table 10. Extrapolated thickness of the transformed 
layer in the grooves at the origin of rifling of caliber .50 
erosion-testing gun. 


Powder 

Flame 

temper¬ 

ature 

(K) 

Transformed layer thickness 

PE 

bullets* 

(in.) 

AT 

bullets f 
(in.) 

High level 
tests 
(in.) 

RDX-5060 

2320 

0.0010 



Cordite N 

2469 

0.0006 

0.0008 


FNH/P 

2480 


0.0008 


FNH-M1 

2483 

0.0009 

0.0008 

0.00085 

RDX-5061 

2560 

0.0013 



NH-M1 

2807 

0.0011 

0.0012 


Pyro 

2814 

0.0012 

0.0012 


Cordite NQ 

2864 

0.0012 

0.0013 


IMR 

2938 

0.0015 

0.0015 

0.0015 

RDX-5059 

3080 


0.0011 


FNH-M5 

3268 

0.0019 

0.0019 


FNH-M2 

3564 

0.0021 

0.0021 

0.0022 

40% NG 

3945 



0.0023 

60% NG 

4300 



0.0024 


* PE bullets are pre-engraved. (See Section 15.2.1.) 
t AT bullets are artillery-type. (See Section 15.2.1.) 


were studied at the end of 1, 5, 10, and 70 rounds in 
order to find the amounts of transformed layer which 
had formed at these stages. The results showed that 
the layer had formed during the firing of the first 
round and that, although very nearly the full thick¬ 
ness was formed at the origin of rifling, somewhat less 
than full thickness was formed toward the muzzle; 
thus the distribution of the layer changed somewhat 
during the firing of subsequent rounds as will be seen 
in Figure 8, wherein distribution curves are plotted 
for 1, 5, 10, and 70 rounds fired with pre-engraved 
bullets and double-base powder containing 40% 
nitroglycerin. 

Full data on the rate of formation of the trans¬ 
formed layer may be obtained from Table 11. Anal¬ 
ysis of these data demonstrates the extreme rapidity 
of the formation of the layer and shows that the re¬ 
gion in the barrel throughout which the layer thick¬ 
ness is independent of the number of rounds fired 
lengthens as the flame temperature of the powder 
increases. It must be noted that general melting is so 
pronounced during the firing of the very hot 40% and 
60% NG powders, that appreciable material is re¬ 
moved from the surface at the same time that the 
layer is altering. Accordingly, and as a first approxi¬ 
mation, observed layer thicknesses in Table 10 have 
been corrected by the addition of the observed 
change in groove radius during the firing of the last 
round. The correction is entirely negligible in the 
case of other powders. 


CONFIDENTIAL 


























TESTS IN CALIBER .50 EROSION-TESTING GUN 


321 


Table 11. Transformed layer thickness as function of 
rounds fired at 1, 2, and 4 in. from the origin of 
rifling in caliber .50 erosion-testing gun. 


Rounds Transformed layer thickness at 


Powder 

fired 

\ in. 

1 in. 

2 in. 

4 in. 

FNH-M1 

1 

0.0003 + 

0.0002 

0.0001 + 

0 


5 

0.0004 

0.00025 

0.00015 

0 


10 

0.0005 + 

0.0004 

0.0003 - 

0.0001 - 


70 

0.0008 

0.00075 

0.0006 

0.00025 

IMR 

1 

0.0008 

0.0007 

0.0006 

0.0005 


5 

0.0009 + 

0.0009 

0.0008 + 

0.0008- 


10 

0.0010 

0.0010- 

0.0009 

0.0008 


70 

0.0012 

0.0011- 

0.0010 

0.0009 

FNH-M2 

1 

0.0020 

0.0018 

0.0014 

0.0012 


5 

0.0020 

0.0018 

0.0015 

0.0013 


10 

0.0020 

0.0018 

0.0017 

0.0015 


70 

0.0022 

0.0021 

0.0020 

0.0019 

40% NG 

1 

0.0021 

0.00195 

0.0018 

0.0013 


5 

0.0021 

0.00195 

* 0.0019 

0.0015 


10 

0.0021 

0.00195 

0.0019 

0.0016 


70 

0.0021 

0.00195 

0.0019 

0.0017 

60% NG 

1 

0.0021 

0.0019 + 

0.0017 

0.0013- 


5 

0.0021 

0.0019 + 

0.0018 

0.0016- 


10 

0.0021 + 

0.0020 

0.0019 

0.0018- 


70 



0.0020 

0.0020 


Relation to Erosion Mechanism. It follows from the 
fact that a continuous, thick transformed layer forms 
during the firing of a single round that erosion of a 
gun steel barrel is really the erosion of a transformed 
layer. Thus a full knowledge of the nature of the 
transformed layer is essential if the mechanism of 
such erosion is to be discovered. The nature of the 
layer has been discussed in Sections 12.1.2 and 13.2.3. 
It remains only to reemphasize that the layer defines 
the region which is rendered austenitic during the 
firing of any single round and that there is consider¬ 
able evidence that the hot, austenitic layer has a high 
chemical reactivity with the constituents of the pow¬ 
der gases. 


15 3 4 Chemical Products on the Bore Surface 

The eroded surface of each fired barrel was sub¬ 
jected to x-ray diffraction analysis with the purpose 
of finding any differences in chemical nature of the 
surfaces resulting from the firing of different powders, 
which is discussed at length in Section 12.2. The 
diffraction patterns were obtained by allowing mo¬ 
lybdenum Ka radiation to strike the eroded surfaces 
at small angles. 

In addition to ferrite (alpha-iron modification) 
which was found on each film, four other constituents 
were found in varying amounts: austenite (gamma- 
iron modification), cementite (Fe 3 C), wiistite (FeO) 
and the epsilon phase of iron nitride (FC 2 N*). It was 
noted that the diffraction lines of the ferrite were 
often quite broad, particularly in the barrels fired 
with the double base powders which have the higher 
flame temperatures. This broadening may actually 
indicate the presence of distorted-cubic martensite 
formed by quenching austenite. Austenite itself is 
never retained by quenching alone, but only if the 
gamma-iron takes up in solid solution a considerable 
amount (up to about 0.6%) of carbon or nitrogen or 
both. Accordingly, the austenite found was probably 
enriched in carbon or nitrogen or both. 

A purely qualitative analysis of the various eroded 
surfaces is given in Table 12. The designated amounts 
are only estimates. Evidently, there is an absorption 
of carbon or nitrogen or both and stabilization, in a 
superficial layer, of austenite regardless of the powder 
fired. 

In addition, two tendencies may be noted: a tend¬ 
ency toward carburizing and nitriding with the 
formation of cementite and iron nitride; and a ten¬ 
dency toward oxidizing with the formation of wiistite. 


Table 12. Qualitative x-ray diffraction analysis of eroded bore surfaces of erosion-testing gun barrels fired with different 
powders. 


Flame 

Powder temperature 

(K) 

Austenite 

(r-Fe) 

Cementite 

(Fe 3 C) 

Wiistite 

(FeO) 

Epsilon 

iron 

nitride 

(Fe 2 N x ) 

Miscellaneous 

RDX-5060 

2320 

Present 

Present 

Possibly 

Present 

Pb + X 

Cordite N 

2469 

Large 

Small 

Small 

Large 

Pb + X 

FNH-M1 

2483 

Large 

Small 

Probably 

Large 

Pb + X 

RDX-5061 

2560 

Present 

Probably 

Present 

Present 

Pb + X 

NH-M1 

2807 

Large 

Present 

Probably 

Probably 

Pb 

Pyro 

2814 

Large 

None 

Small 

Large 


Cordite NQ 

2864 

Large 

None 

Small 

Present 

Pb 

IMR 

2938 

Large 

Present 

None 

None 

Pb 

FNH-M5 

3268 

Large 

None 

Large 

None 


FNH-M2 

3564 

Large 

None 

Large 

None 



CONFIDENTIAL 













322 


EROSION OF GUN STEEL BY DIFFERENT PROPELLANTS 


High flame temperatures appear to encourage the 
latter reaction and discourage the former; while low 
temperatures seem to permit both, although oxidiza¬ 
tion is not very pronounced. 

Metallic lead (Pb) was found on many of the sur¬ 
faces. Possibly the lead came from lead thiocyanate 
in the primers and was volatilized in the tests with 
very hot powders. 

Unidentified substances (X) were found on several 
of the surfaces. 


i5.4 VENT PLUG TESTS 

The erosion produced in vent plugs (Section 15.2.2) 
was measured in terms of the weight loss of each 
plug tested at the levels of severity afforded by differ¬ 
ent vent sizes. The measurements were usually made 
after the firing of 5, 10, and 15 rounds with each of 
the different powders. Most powders were tested 
with four different vent sizes: Xe, %2, and Xq in., 
but the RDX powders were tested only with three: 

}/g, and Xo in. NH-M1 powder (for 37-mm gun) 
was tested along with the latter to provide a reference 
for the comparison of all powders. 

A summary of the results in the form of averages of 
the two runs after 15 rounds are listed in the two 
parts of Table 13 for all of the powders tested. Ero- 


Table 13. Erosion of vents of different sizes expressed 
in terms of average weight loss after 15 rounds. 


Powder 

Average weight loss (mg) 
/4-in. Brin. 54- in. 

Be-in. 

FNH/P 

0 

14.3 

92.2 

212.8 

FNH-M1 

0 

23.6 

127.8 

246.2 

NH-M 1-3727 

0 

38.1 

120.1 

230.8 

NH-M 1-4974 

0 

12.2 

108.6 

210.2 

Pyro 

1.9 

18.3 

65.5 

212.1 

IMR 

2.8 

35.6 

118.1 

263.7 

FNH-M2 

24.8 

339.9 

512.4 

1069.9 


54-in. 

Vs- in. 


Be-in. 

RDX-6079 

3.3 

34.5 


276.6 

RDX-6080 

7.4 

67.4 


330.6 

NH-M 1-3727 

2.7 

30.5 


285.0 

RDX-6081 

4.1 

59.7 


317.6 

RDX-6082 

6.8 

67.3 


386.1 

RDX-6083 

18.5 

118.1 


513.3 


sion was only very small in the largest vents. Hence 
tests with large vents do not distinguish among the 
less erosive powders. The more severe conditions in 
the smaller vents permitted a finer classification. 


15-41 Relative Erosiveness of Powders 

The powders tested in 3^8-in. and 3d6-in. vents (the 
only sizes used for both the RDX and the other 
powders) are listed in Table 14 according to their 


Table 14. Relative erosiveness of powders tested in 
Bi-in. and 14-in. vent plugs. 


Brin, vent 

Brin, vent 


Degree of 


Degree of 

Powder 

erosiveness 

Powder 

erosiveness 

NH-M 1-4974 

0.32 

NH-M1-4974 

0.91 

FNH/P 

0.38 

Pyro 1 

0.92 

Pyro 

0.48 

FNH/PJ 

FNH-M1 

0.62 

RDX-6079 

0.97 

IMR 

0.94 

NH-M1-3727 

1.00 

NH-M1-3727 

1.00 

FNH-M1 

1.06 

RDX-6079 

1.13 

RDX-6081 

1.12 

RDX-6081 

1.96 

IMR 

1.14 

RDX-6080) 

2.21 

RDX-6080 

1.16 

RDX-6082/ 

RDX-6082 

1.35 

RDX-6083 

3.87 

RDX-6083 

1.80 

FNH-M2 

8.92 

FNH-M2 

4.62 


relative order of erosiveness. In both lists an erosive¬ 
ness of unity is assigned to NH-M1 powder, lot No. 
3727. The numbers used to differentiate the powders 
of the same types are experimental lot numbers ex¬ 
cept in the case of NH-M1-3727 which is a standard 
lot of powder for 37-mm gun, M1916. 

There is not a one-to-one correspondence between 
these two orders but in many significant features 
they are quite similar. By consideration of the partial 
series determined with the other vent sizes and weight¬ 
ing the results slightly toward low severities so as to 
approximate better the actual erosive loss found in 
real guns, it is possible to organize an overall 
erosiveness series by groups : 

NH-M 1-4974, FNH/P, Pyro 
< NH-M 1-3727, RDX-6079 (?) 

< FNH-M1 < IMR < RDX-6081 
< RDX-6080, RDX-6082 
< RDX-6083 < FNH-M2 

This series differs in a number of respects from that 
determined by firings in the caliber .50 erosion-testing 
gun, as given at the end of Section 15.3.1. It will be 
noted from a comparison of Tables 2 and 4 that in 
most cases where RDX powders and single-base pow¬ 
ders have nearly the same flame temperature, (that 
is, RDX-6081, and NH-M 1-3727) the former are 
considerably more erosive than the latter. This fact 
emphasizes the very great importance of chemical 


CONFIDENTIAL 














CORRELATION OF EROSION AND THERMAL FACTORS 


323 


action by the powder gases in the problem of erosion, 
for in such cases thermal conditions are closely sim¬ 
ilar and the only important difference is to be found 
in the chemical composition of the powder gases. It 
follows that the overall correlations of erosive effects 
with temperature which are made below are valid 
only so long as the chemical compositions of the gases 
are roughly similar, for example, as they appear to be 
among the gases from single-base and double-base 
powders. 


The position of RDX-6079 is somewhat uncertain 
at best. Its present position is only one group re¬ 
moved from that of the matching RDX-5060. The 
sole other discrepancy is the position of FNH-Ml. 
There seems to be no explanation for the lack of 
agreement in its case. 

In spite of these anomalies it may be concluded 
that vent plug tests of powders are useful in giving a 
rough idea of relative erosiveness. 


15 4 2 Comparison of Erosiveness 
in Guns and Vents 

The significance of these tests is clarified when 
comparison is made with the erosion rates found by 
firing the erosion-testing gun (Section 15.3). In Table 
15 the various powders are listed in order of increas- 


Table 15. Order of increasing powder erosiveness as 
found by tests in the vent plug and in the caliber .50 
erosion-testing gun. 


*- 

Vent plug 

Gun 

/FNH/P 

(RDX-5060 

1 Pyro 

(FNH/P 

JNH-M1 

(FNH-Ml 

\ RDX-6079 

NH-M1 

FNH-Ml 

Pyro 

IMR 

IMR 

FNH-M2 

FNH-M2 


ing erosiveness as determined in the vent plug and in 
the pre-engraved bullet series in the erosion testing 
gun. Only the powders tested by both methods are 
included, except for powders RDX-6079 and RDX- 
5060, which were nearly the same in composition and 
flame temperature. 

Some rationalization of differences between the 
two listings is possible. For example, the largest dif¬ 
ference is the position of Pyro powder, which was 
classified in the least erosive group by the vent-plug 
tests and in an intermediate position by the gun tests. 
It will be noted that in tests with the largest vent in 
which erosion is most nearly equivalent in amount to 
that found in the gun, Pyro was ranked between 
IMR and the other less erosive powders as it is in the 
gun listing. It seems possible that had those low- 
severity tests been able to distinguish among the less 
erosive powders it might have led to a listing more 
like that found with the gun and would certainly 
have brought agreement with respect to Pyro. 


15 5 CORRELATION OF EROSION, THERMAL 
TRANSFORMATION, HEAT INPUT, AND 
FLAME TEMPERATURE 

15-51 Introduction 

Throughout this discussion relationships between 
the rate of erosion as measured by increase in bore 
diameter, the thermal transformation of the barrel 
steel, the heat input to the bore surface and the flame 
temperature of the powder have been implied. A de¬ 
scription of these relationships in definite form pro¬ 
vides a good summary of the foregoing sections as 
well as a satisfying demonstration of the intimate 
connection among these various aspects of gun ero¬ 
sion. 

Temperature may be regarded as the key which 
opens the door to understanding of the nature of gun 
barrel erosion and in particular to an understanding 
of erosion as caused by different propellants. The 
temperature of the bore surface determines the me¬ 
chanical properties of that surface, the kinds and the 
rates of chemical reactions which will occur on and in 
that surface, and the possibility of liquefaction of 
that surface or of the chemical reaction products that 
form there. The sum of these effects during firing, 
discussed in more detail in Chapter 13, embodies the 
mechanism of gun erosion. 

Any correlation of erosive effects with temperature 
should properly be made with respect to a bore- 
surface temperature, but such a temperature is not 
easy to define. First of all, the temperature of the 
immediate bore surface fluctuates with extreme ra¬ 
pidity and since erosive effects are not instantaneous, 
some suitable time-average needs to be chosen. Sec¬ 
ondly, erosive effects are not two-dimensional but 
occur within small, albeit, finite volumes. Tempera¬ 
ture distribution falls off rapidly inward from the 
bore surface, so that some suitable space-average 
needs to be chosen for the reference temperature. 


CONFIDENTIAL 









324 


EROSION OF GUN STEEL BY DIFFERENT PROPELLANTS 


Problems of this sort are not impossible to solve 
and suitable methods of calculating these average or 
“effective” temperatures have been worked out. 48 
They are, nevertheless, rather laborious. 

A simpler approach is to find some easily measured 
quantity which is related to the bore surface temper¬ 
ature. The adiabatic flame temperature of the pow¬ 
der (Section 2.4.3) is perhaps the most convenient 
such quantity, although the relation is not a simple 
one, involving as it does the intermediate relations of 
gas temperature and heat input. Reliance upon it as 
a parameter measuring the bore surface temperature 
is greatly strengthened by the existence of a simple 
linear relationship between it and the actual meas¬ 
ured heat input to the bore which will be demon¬ 
strated in Section 15.5.2. It is not unreasonable, then, 
to attempt to correlate some of the aspects of meas¬ 
ured erosion of these barrels against the flame tem¬ 
perature of the powder fired. 

15 5 2 Dimensional Erosion Rate versus 
Flame Temperature 

The dimensional erosion rate is, of course, not a 
function of the flame temperature alone, but depends 
also on the type of bullets fired and on the loading 


conditions. The type of bullet is a truly independent 
factor; the loading condition is probably not truly 
independent. That is to say, if the loading condition 
is adjusted to give a higher ballistic level with the 
same powder, the relation between flame tempera¬ 
ture and bore-surface temperature will change so as 
to raise the latter. The erosion rate and attendant 
phenomena are therefore increased in magnitude as 
will be illustrated below. The dimensional erosion 
rate may be studied with reference to flame tempera¬ 
ture, therefore, only if other factors are kept constant 
or if their influence is quantitatively understood. 

The data of Table 8 present the relation between 
erosion rate and flame temperature for the pre-en- 
graved bullet, the artillery-type bullet, and the high 
ballistic level series. These data are plotted in Figure 
9 where the curve for each series is seen to be similar 
in shape to an exponential curve. An exponential re¬ 
lationship is strongly implied by the straight lines 
that result if the same data are plotted on semilog- 
arithmic paper. 

For a given change in flame temperature, the ero¬ 
sion rate appears to increase more rapidly at higher 
flame temperature. Evidently also the use either of 
engraving type of bullets or of a higher ballistic level 
increases the erosion rate at all temperatures. 



2400 2600 2800 3000 3200 3400 3600 


POWDER FLAME TEMPERATURE IN DEGREES K 



2800 3400 3800 4200 

FLAME TEMPERATURE IN DEGREES K 


Figure 9. Erosion rate (in units of 0.00001 in. per round) versus flame temperature for the pre-engraved bullet, the 
artillery bullet, and the high ballistic level series. (This figure has appeared as Figures 18 and 20 in NDRC Report 
No. A-451.) 


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MODIFICATION OF PROPELLANTS 


325 


15 5 3 Heat Input versus Flame Temperature 

The data on heat input reported in Table 9 are 
plotted in Figure 10, where the linear relationship 
between heat input and flame temperature is evident. 
It has been suggested that the deviation from linear¬ 
ity at very high flame temperatures is probably the 
result of absorption of latent heat of fusion during 
the considerable melting of gun steel that appears to 
accompany the firing even of a single round. 



Figure 10. Heat input versus flame temperature. 
(This figure has appeared as Figure 23 in NDRC Re¬ 
port No. A-451.) 


Since the heat input should be more directly asso¬ 
ciated with bore-surface temperature than is the 
flame temperature, it is reasonable to expect that the 
plotting of erosion rate versus heat input should give 
a smoother curve. This deduction is borne out by the 
facts. If the erosion rate is first plotted against heat 
input and then the heat input scale is converted to 
flame temperature by means of the linear relationship 
of Figure 10, the results of Figure 9 are again obtained 
with the difference that there is a better fit of the 
experimental points to the exponential curves showing 
erosion rate versus flame temperature. 

Heat input for a given flame temperature should 
also increase with the ballistic level. The data now 
available do not permit an investigation of this point. 

15.5.4 Thermal Transformation versus 
Flame Temperature 

By virtue of its rapid rate of formation which 
permits it to renew itself during each round, the 
transformed layer presents a characteristic thickness 
at the origin of rifling which is independent of the 
type of bullets or the number of rounds fired. This 
characteristic thickness may be expressed as a func¬ 


tion of the flame temperature. As before, the thick¬ 
ness observed with a given flame temperature should 
increase with the ballistic level, but the available 
data do not permit an investigation of this point. 

The data on thickness of the transformed layer 
which were presented in Table 10 are plotted versus 
flame temperature in Figure 11. A straight line may 
be drawn through points for powders not containing 
RDX, although at very high flame temperatures the 
deviations are greater than at low temperatures. Per¬ 
haps this is caused by the absorption of the latent 
heat of fusion having changed the thermal conditions. 
The distribution of the points for the three RDX 
powders seems to be random. 



Figure 11. Transformed layer thickness versus flame 
temperature. (This figure has appeared as Figure 64 in 
NDRC Report No. A-451.) 


The mechanism of thermal transformation has 
been discussed in Section 13.2.3 and a method of 
calculation has been presented 124 which permits an 
explanation of the results of Figure 11. 

156 MODIFICATION OF PROPELLANTS e 
1561 Introduction 

The foregoing section points out the dependence of 
dimensional erosion rate upon the adiabatic flame 
temperature of the powder. From this it would seem 
that a mitigation of erosion might be obtained by 
developing powders of low flame temperature but 
with high impetus. This was suggested in Section 

e An Ad Hoc Committee on Internal Ballistics appointed by 
the Chairman of the NDRC in 1942, recommended a joint in¬ 
vestigation of this subject by the units of NDRC which were 
subsequently designated as Divisions 1 and 8. 143 


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326 


EROSION OF GUN STEEL BY DIFFERENT PROPELLANTS 


3.5.4 for hypervelocity guns in which erosion with the 
usual propellants is a limiting factor. It should be 
noted, however, that the correlation of erosion with 
flame temperature or heat input holds true only for 
powders with similar composition. The RDX series 
of poAvders f must be considered separately from the 
other powders. In both the studies with vent plugs 
(Section 15.4.1) and with the caliber .50 erosion¬ 
testing gun (Table 8) it was found that RDX pow¬ 
ders though cooler than the older standard powders 
are more erosive. Thus, while it is clear that the tem¬ 
perature of the gases resulting from the combustion 
of a propellant is probably the most important factor 
to be considered in the development of less erosive 
powders, the composition of the gases is also impor¬ 
tant, and the discussion which follows is warranted. 

15,6,2 Variations in Types of 

Erosion Products 

From studies of eroded guns and from the results 
of laboratory firing tests a great deal has been learned 
about the effects on gun bores of the various com¬ 
ponents of the powder gases from different propel¬ 
lants. X-ray- and electron-diffraction examination of 
steel gun bores, test blocks, and filings exposed to the 
gases from the usual single- and double-base propel¬ 
lants (Sections 12.2 and 15.3.4) has proved that the 
nitrocellulose powders (single-base) are carburizing 
whereas those containing nitroglycerin in addition to 
nitrocellulose (double-base) are oxidizing. Further 
evidence of this fact was obtained in the studies of 
carbon penetration (Section 14.2.5). 

Metallographic and chemical studies of the surface 
layers in eroded guns (Chapter 12) have indicated 
that the dominant erosion process in guns fired with 
single-base powders is the removal of partially lique¬ 
fied material high in carbon and nitrogen, which has 
resulted from the chemical alteration of the steel by 
the powder gases. This material has a fusion range 
about 300 C lower than the melting point of gun 
steel. On the other hand, liquefaction in guns in 
which oxidation by the gases from powders contain¬ 
ing nitroglycerin takes place is that of unaltered or 
perhaps only slightly altered steel, as is brought out 
in Section 13.2.4. 

The obvious conclusion from the evidence sum¬ 
marized above is that a less erosive powder might be 

f These powders had been investigated by Division 8, 
NDRC. They were also among those investigated in Cana¬ 
da 371 during World War II. 


one that was only a slight modification of the stan¬ 
dard nitrocellulose-nitroglycerin type of propellants. 
It should definitely be less carburizing than the 
single-base powders so that low-melting products 
would not be formed on bore surfaces but should not 
have as high an adiabatic flame temperature as the 
usual double-base powders which cause melting of 
the steel itself. Although oxidation does not promote 
liquefaction to the same extent as carburization, less 
oxidation than is obtained with the usual, relatively 
erosive double-base powders would certainly be de¬ 
sirable, since oxide films are presumably readily re¬ 
moved by mechanical forces during firing, and would 
probably be obtained in the case of a powder that 
was developed to have a lower flame temperature. 

Another modification of powders is suggested in 
Section 14.4.8 which points out the desirability of 
decreasing the amount of sulfur in charges. 


15,6,3 Thermodynamic Considerations 
Single-Base and Double-Base Powders 

The same thermodynamic considerations that led 
to an evaluation of the chemical factors in the causes 
of erosion, discussed in Section 13.3, may be applied 
to the problem of finding a less erosive propellant. In 
the reactions of the powder gases with gun steel the 
most important components of this gas are carbon 
monoxide and carbon dioxide. Whether oxidation or 
carburization of the bore surface occurs depends on 
the CO/C0 2 ratio in the gases at the effective temper¬ 
ature, for this ratio decreases with the temperature 
of the gases. Powders containing nitroglycerin have a 
higher oxygen balance than those containing only 
nitrocellulose, thus the C0/C0 2 ratios of the former 
are lower than those of the latter (Figure 9 of Chap¬ 
ter 2) and the gases from double-base powders may 
be considered more oxidizing or less carburizing. Fig¬ 
ures 6 to 8 of Chapter 13 show fields for the ultimate 
or major products resulting from the interaction of 
steel and powder gases. In order to obtain boundary 
conditions so that oxidation and carburization were 
equal or perhaps even mutually suppressed, an in¬ 
crease in flame temperature and an appreciable de¬ 
crease in the CO/C0 2 ratio of the gases from a single¬ 
base powder would be required. It was suggested 138 
that a nitrocellulose powder to which a small amount 
of nitroglycerin (perhaps 1%) had been added should 
be investigated. 


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MODIFICATION OF PROPELLANTS 


327 


RDX Powders 

According to the above line of reasoning, RDX 
powders might be expected to be more erosive than 
single-base powders of the same flame temperature, 
which actually is the case. Figure 9 of Chapter 2 
shows that the former, because of their higher 
C0/C0 2 ratios, should be much more carburizing 
than the latter. The erosivity of these newer propel¬ 
lants, however, does not bear any obvious relation¬ 
ship to their high CO/CO 2 ratios, for studies of car¬ 
bon penetration with an RDX powder (Section 
14.2.5) indicated that the reaction products were not 
as high in carbon as those obtained from the reaction 
of a single-base powder at the same bore-surface 
temperature. It may be that in the case of the RDX 
powders other reactions besides those of the carbon 
gases with the steel assume more importance than 
they do in the case of the nitrocellulose propellants. 
Indications that this may be true have been found in 
the few instances in which erosion products resulting 
from the use of RDX powders have been examined 
by diffraction techniques. 137 

15 6 4 Experiments with 

Powders Containing Ferrosilicon 

Erosion might be diminished by maintaining a 
sufficiently high temperature to avoid carburization 
and nitriding but at the same time preventing oxida¬ 
tion by the addition of a deoxidizer. Gun steel filings, 
mixed with a charge of double-base powder to which 
ferrosilicon had also been added, were fired in a cal¬ 
iber .30 rifle according to the technique described in 
Section 11.2.6. 49 No ferrous oxide (FeO) was identi¬ 
fied in the collected products. This compound was 
always found in these experiments with steel or iron 
filings when no deoxidizer had been added to the 


charge. Because of the successful reduction of the 
oxidation in this case, a firing test was made in the 
caliber .50 erosion-testing gun (Section 11.2.1) with 
a 20% nitroglycerin powder to which 5% of ferro¬ 
silicon powder had been added. The incomplete test 49 
showed a slight reduction of the erosion, particularly 
in the grooves. Difficulties encountered in deposition 
of unused ferrosilicon in the grooves and excessive 
muzzle blast caused the test to be abandoned. 

15 6 ^Experiments with Powders Containing 
Low Percentages of Nitroglycerin 

As was pointed out in Section 15.6.3, a nitrocellu¬ 
lose powder to which a small amount of nitroglycerin 
had been added might be less erosive than one con¬ 
taining nitrocellulose alone. It was planned to test 
mixtures of a standard single-base and a standard 
double-base propellant in such proportions that low 
percentages of nitroglycerin would be obtained. Such 
mixtures were to be tested first in erosion vent plugs 
and then in caliber .50 machine gun barrels. 

One mixture containing 5% nitroglycerin was test¬ 
ed in the vent-plug apparatus. Because the effect to 
be studied was essentially chemical in nature, the 
3^-in. vent, which gives reproducible erosion at low 
severity, was used. This mixture was more erosive 
than single-base powder. The erosivity of mixtures 
containing less than 5% nitroglycerin, however, may 
be worth investigating. The original suggestion, it 
should be noted, was that only about 1% nitroglycerin 
should be added. Because of the termination of the 
Division’s activities, further testing was not contin¬ 
ued. Only the preliminary control experiments were 
carried out for the machine gun tests. 110 A continua¬ 
tion of these experiments would be worthwhile; nega¬ 
tive results would be of as much value in this case as 
positive ones. 


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PART V 

EROSION RESISTANT MATERIALS 



Fire is the test of gold; adversity of strong men. 

—Seneca 
“De Providentia” 


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Chapter 16 

SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 

By J. F . Schairen 


161 EROSION RESISTANT MATERIALS 
PROGRAM 

I n an early stage of the studies of gun erosion 
at the Geophysical Laboratory for Division 1, 
NDRC, the conclusion was reached that, because of 
their lack of resistance to thermal and chemical at¬ 
tack by powder gases during firing, no steels or high- 
iron alloys showed promise as bore-surface materials 
under severe firing conditions using conventional pro¬ 
pellants. Yet steels are the only materials of adequate 
strength and ductility that are available in sufficient 
quantities for gun tubes. Therefore, in order to pro¬ 
tect the bore surface of such steel tubes from contact 
with powder gases (at least near the breech end 
where powder gas erosion is most severe), attention 
was concentrated on the development of suitable 
erosion resistant liners, linings, electroplates, and 
other coatings. 

Laboratory tests showed that of all the 92 elements 
then recognized, as shown in Figure 1, only the fol¬ 
lowing pure metals b were resistant to chemical attack 
by the powder gases: chromium, molybdenum, tung¬ 
sten, tantalum, nickel, cobalt, and copper. Only the 
first four of these have a sufficiently high melting 
point for severe service under hypervelocity condi¬ 
tions, where melting is an important factor in the 
failure of a steel gun bore surface. Subsequent tests 
on pure nickel and high-nickel alloys under less severe 
laboratory test conditions and firing tests as gun 
liners showed that these (except nickel-chromium 
alloys with more than 10% chromium) are subject 
to severe intergranular attack by powder gases and 
gave a performance barely equal to or inferior to gun- 
steel. Copper is quite unsuitable for a gun bore sur¬ 
face material because of its softness and low strength 
in addition to its low melting point. Other tests 
showed that in addition to suitable resistance to 
thermal and chemical attack a bore surface material 


a Special Assistant, Division 1, NDRC. (Present address: 
Geophysical Laboratory, Carnegie Institution of Washington.) 

b There is some evidence (Section 16.4.14) that certain of 
the metals of the platinum group may be erosion resistant. On 
account of their scarcity these metals were not seriously con¬ 
sidered in the Division 1 program. 


must have sufficient hardness and strength at tem¬ 
peratures attained during firing to prevent deforma¬ 
tion of the rifling by impact of the projectile and 
must be sufficiently ductile to prevent serious failure 
by cracking. 

In the fall of 1942 efforts were started by Division 
1 on the preparation of chromium and molybdenum 
in form suitable for use as gun liners. By the following 
summer preliminary tests of molybdenum liners had 
emphasized the importance of hot hardness as a char¬ 
acteristic of a successful gun liner material. Further 
study of this phase of the subject led to the discovery 
that the stellites (Chapter 19), which are cobalt- 
chromium alloys that have the property of hot-hard¬ 
ness, are erosion resistant as long as the bore-surface 
temperature is not too high. 

By that time the experience of aerial combat during 
World War II had indicated that erosion was limiting 
the performance of the caliber .50 aircraft machine 
gun. Application of the discovery of the erosion resis¬ 
tance of stellites to this problem led to a remarkable 
increase in the performance level of this gun (Chapter 
22). A parallel attack on this same problem led to the 
development of nitrided, chromium-plated caliber .50 
barrels (Chapter 23). Eventually, it was found that 
an even better barrel was obtained by using a stellite 
liner with the steel bore chromium-plated ahead of it 
in such a manner as to constrict the bore at the muz¬ 
zle (Chapter 24). Furthermore it was found that pro¬ 
vided the steel barrel was strengthened by making it 
slightly heavier (especially at the forward end of the 
liner) and perhaps by using a special steel having 
greater strength at high temperatures even better 
performance could be attained (Chapter 24). 

Experience with stellite liners in the caliber .60 
machine gun, which has a muzzle velocity of slightly 
over 3,500 fps, showed that this alloy is “marginal” 
with respect to its use in a hypervelocity gun. In this 
particular application a stellite liner lasts long enough 
to furnish a useful gun barrel life, but when it fails, 
it does so by melting along surface cracks. That fact 
coupled with the observation that the surface of a 
stellite liner melts when fired with double-base pow¬ 
der, even at velocities around 3,000 fps, showed that 
a material of higher melting point was needed for 


CONFIDENTIAL 


331 



332 


SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 



H 








He 

Li 

Be 

B 

C 

N 

O 

F 


Ne 

Na 

Mg 

Al 

Si 

P 

S 

Cl 


A 

K 

Ca 

Sc 

Ti 

V 

Cr 

Mn 

Fe,Co,Ni 


Cu 

Zn 

Ga 

Ge 

As 

Se 

Br 


Kr 

Rb 

Sr 

Y 

Zr 

Cb 

Mo 

Ma 

Ru, Rh, Pd 


Ag 

Cd 

In 

Sn 

Sb 

Te 

I 


Xe 

Cs 

Ba 

La 

Ce 









Hf 

Ta 

w 

Re 

Os, Ir, Pt 


Am 

Hg 

Tl 

Pb 

Bi 

Po 

Ab 


Rn 

Vi 

Ra 

Ac 

Th 

Pa 

U 






















Figure 1. Only a few of the 92 elements in the periodic table have been found to be resistant to chemical and thermal 
attack by the powder gases. The symbols of the rare earth elements have been omitted from this table. 


general use in hypervelocity guns, that is, ones fired 
at muzzle velocities greater than 3,500 fps. Hence the 
search for such a material was continued at the same 
time that further efforts were made to extend the 
application of stellite. 

During World War II, caliber .50 gun barrels that 
had been nitrided and chromium plated and others 
that had had stellite liners inserted in them were used 
in combat. Production of stellite-lined barrels of other 
sizes was ready to start at the time of Japanese sur¬ 
render. The further application of stellite and other 
hot-hard alloys to small arms barrels was continued 
by the Crane Company for the War Department. 
Continuation of the investigation of chromium elec¬ 
troplates, of duplex electroplates of chromium and 
other metals, and of alloy electroplates of various 
pairs of metals (Chapter 20) at the National Bureau 
of Standards was supported jointly by the War and 
Navy Departments. 

The most promising material for service in hyper¬ 
velocity guns appears to be a hardened molybdenum 
(Chapter 18). Sufficient progress was made in its 
development during the war so that the Navy De¬ 
partment continued to support the efforts of the 
Westinghouse Electric Company to make this ma¬ 
terial in a form suitable for large gun liners, following 


the plans described in Section 33.1.3. Similarly, the War 
Department, under contract with the Union Carbide 
and Carbon Research Laboratories, continued the 
development of chromium-base alloys, which also 
appear very promising for hypervelocity service 
(Chapter 17). Vapor-phase plating (Chapter 21) does 
not appear to be suitable for gun bores, and therefore, 
its development would not have been continued by 
Division 1 for this purpose, although it may have 
industrial applications. 

Thus the resistant-materials program of Division 
1, NDRC, during the course of 33^ years led to the 
development of a very successful solution to the ero¬ 
sion problem in machine guns and narrowed the 
search for bore-surface materials capable of outstand¬ 
ing performance under hypervelocity conditions to 
three clearly-defined programs, all of which were 
subsequently pursued by the Armed Services. The 
present chapter surveys the whole resistant materials 
program. 0 In so doing, the successful developments, 
which receive more extended treatment in subsequent 
chapters (already noted in the foregoing paragraphs), 
are summarized in order to give proper perspective 

c The work carried out by the contractors of Division 1, 
NDRC, in the development of erosion-resistant materials is 
described in a series of reports listed in the Bibliography. 


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PROPERTIES OF AN IDEAL EROSION RESISTANT MATERIAL 


333 


to the large number of unsuccessful tests of other 
materials. 

16 2 PROPERTIES OF AN IDEAL EROSION 
RESISTANT MATERIAL 

16-21 Introduction 

Some very definite ideas on the properties of an 
ideal erosion resistant material have evolved during 
the course of Division l’s investigation of gun erosion. 
These concepts are the result of an analysis of the 
results of laboratory tests of materials and firing tests 
of gun barrels with and without liners, platings or 
coatings of special metals or alloys (described in the 
following sections of this chapter and the remainder 
of Part V), combined with an analysis and interpreta¬ 
tion of the results of ballistic studies, observations on 
eroded gun bores, and studies of the mechanism of 
erosion (described in Parts II, III, and IV). 

The properties required of a satisfactory bore-sur¬ 
face material depend on the conditions under which 
it is to be used. The suitability of a given material 
for use as a bore surface in a particular gun tube is 
determined by the characteristics required or desired 
for that weapon with respect to size, rate of fire, 
muzzle velocity, type of ammunition (design of pro¬ 
jectile and choice of propellant), the pressure-tem¬ 
perature-time characteristics of the powder gas, firing 
schedule, and minimum useful life. Conversely, the 
ability to attain some of the desired gun character¬ 
istics will depend upon obtaining a suitable bore- 
surface material. 

The prime requisites of a satisfactory bore-surface 
material for any gun tube are resistance to thermal 
and chemical attack by the powder gas and suitable 
mechanical properties. 

16-2-2 Thermal Resistance d 

Resistance to thermal attack involves four separate 
requirements. First, the material must have a com¬ 
bination of high melting point, high specific heat, and 
high thermal conductivity, such that the maximum 
temperature attained by the bore surface (Section 
5.4) will always be well below the melting point of the 
material. For hypervelocity guns, the melting point 
of the bore surface material should be at least 1500 C 
and preferably very much higher. 

d See Section 13.2 for a discussion of the thermal factors in 
the erosion of steel guns. 


The second requirement is a high resistance to 
thermal and chemical shock (Section 13.1.2) as evi¬ 
denced by a minimum tendency for the surface to 
crack under a rapid heating and cooling cycle in the 
presence of chemically active powder gases. 

In the third place, the material must undergo no 
abrupt volume changes as a result of the cyclic 
changes of temperature and pressure during the firing. 
In the case of a coating on steel such volume changes 
may cause poor adhesion and cracking and thus per¬ 
mit attack of the underlying steel by the powder 
gases, and in the case of a solid liner they may cause 
either serious cracking or severe constriction of the 
bore. 

As a fourth requirement, the coefficient of thermal 
expansion should be low so that at elevated tempera¬ 
tures the rotating projectile does not lose engage¬ 
ment with the rifling and fail to attain adequate spin. 

16-2-3 Chemical Resistance 6 

Chemical resistance requires either that the ma¬ 
terial be inert to the powder gases at the tempera¬ 
tures attained by the bore surface in contact with the 
hot gases, or that, if chemical reaction does take 
place, a thin adherent protective film be formed. The 
rate of chemical reaction is involved in chemical re¬ 
sistance : the rate must be so low that no deleterious 
attack occurs in the short time that hot gases are in 
contact with hot metal in a particular weapon. 

16-2-4 Mechanical Properties f 

In addition to resistance to thermal and chemical 
attack by hot powder gases, a satisfactory bore-sur¬ 
face material in any weapon must remain in place 
during firing, suffer no large permanent change in 
dimensions, show no permanent deformation or flow- 
age and must not wear excessively or disintegrate by 
brittle failure after impact. The mechanical proper¬ 
ties of the material both at room temperature® and at 
elevated temperatures determine whether a given 
material with thermochemical resistance to attack by 
powder gases can meet these additional mechanical 

e See Section 13.3 for a discussion of the chemical factors in 
the erosion of steel guns. 

f See Section 13.4 for a discussion of the mechanical factors 
in the erosion of steel guns. 

g In gun barrels for use at high altitudes in aircraft or under 
severe arctic conditions, some of the low-temperature (down to 
—100 F) properties of both the barrel steel and special bore 
surface material are also important. 


CONFIDENTIAL 






334 


SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 


requirements of a satisfactory bore-surface material. 

The strength at elevated temperatures must be 
high. If engraving-type projectiles are used, this 
strength is necessary to prevent permanent deforma¬ 
tion (swaging) of the rifling by impact of the projec¬ 
tile; however, this requirement can be made less 
stringent by the use of pre-engraved projectiles (Sec¬ 
tion 27.3 and Chapter 31). Resistance to swaging, 
although it is related to hot-hardness and hot- 
strength, is not determined by these properties alone, 
but also depends on high-velocity impact resistance, 
which is difficult to measure. 

Resistance to abrasion and wear at high tempera¬ 
tures, which presumably is another corollary of high 
hot-hardness, must also be high. The ductility of the 
material and its impact strength both hot and cold, 
must be great enough to prevent cracking and brittle 
failure during the severe thermal cycle of repeated 
firing. 

Because most erosion resistant materials are applied 
as a liner, a lining, or as a bore coating in a gun-steel 
barrel, it is preferable that the thermal and elastic 
properties of the material be close to those of gun 
steel. 

16,2-5 Other Properties 

In addition to the prime requisites—thermochemi¬ 
cal resistance and suitable mechanical properties— 
just discussed, other desirable properties of a suitable 
bore-surface material are availability in the quanti¬ 
ties desired and ease of application. In order to per¬ 
mit its application in the form of a liner or lining, a 
substance must be amenable to fabrication in suitable 
sizes and shapes and must be machineable for boring 
and rifling or it must be applicable as a relatively 
thin bore coating. 

In the next section of this chapter, the test meth¬ 
ods used are described and the results are summar¬ 
ized. Although valuable preliminary information can 
be gained from laboratory tests of materials and the 
choice of materials can thereby be narrowed to a few 
promising substances, the ultimate test of an erosion 
resistant material is a firing test in the gun in which 
it is to be used or under conditions as close as possible 
to this ultimate objective. The specific applications 
of the newly-acquired knowledge of erosion resistant 
materials and erosion processes to improvement in 
the performance of Service weapons and to the design 
and development of new weapons are described in 
Parts VI and VII. 


16.3 TEST methods for evaluating 
PROSPECTIVE MATERIALS 

16,31 Laboratory Explosion Vessels 

Important quantitative and qualitative data on 
the resistance of materials to the erosive action of hot 
powder gases were obtained by the use of laboratory 
explosion vessels. The information gained concerned 
their resistance to melting, chemical attack, and in 
some cases surface cracking, but did not include their 
resistance to the mechanical effects of high pressures 
and deformation or wear by a projectile. 

Erosion Vent Plugs 

The first device used by Division 1 in evaluating 
the resistance of metals and alloys to powder-gas 
erosion was the erosion vent plug. 27 The method of 
erosion vent-testing previously employed was modi¬ 
fied as described in Section 11.2.3. The rate of erosion 
was made less severe and a better control of the rate 
of burning of the powder was achieved by closing the 
muzzle end of the vent plug with a brass rupture disk 
to release the gases at the same predetermined pres¬ 
sure. The amount of metal removed on successive 
firings, as determined by change in weight, combined 
with the results of metallographic examination of the 
surface of the vent hole after firing was completed, 
gave a basis for estimating the erosion resistance of 
the sample. 

In the first tests 27 of nearly 100 metals and alloys 
Y6-in. vents and double-base powder were used. 
Under these relatively severe conditions the amount 
of metal removed by fusion is much greater than that 
lost in any other way, and therefore the results sug¬ 
gest what metals have suitable erosion resistance for 
use under conditions of hypervelocity where bore 
melting may be a most important factor. Only four 
metals (tungsten, tantalum, molybdenum, and chro¬ 
mium) showed promise of a marked improvement 
over gun steels in resistance to thermal and chemical 
attack under hypervelocity conditions while two 
others (copper and nickel) showed a less marked 
improvement. 

Nickel and most nickel alloys (except nickel-chro¬ 
mium alloys with more than 10% chromium) are 
subject to intergranular corrosion when exposed to 
hot powder gases. This tendency was not detected in 
the erosion vent-plug test with 346-in. vents because 
of the removal of metal by fusion. Subsequent vent 
tests at decreased severity described below and tests 


CONFIDENTIAL 



TEST METHODS FOR EVALUATING PROSPECTIVE MATERIALS 


335 


of liners described later (see Sections 16.3.5, 16.3.7, 
16.3.8 and 16.4.9) show the complete unsuitability of 
such materials for a bore surface. 

Subsequent experiments indicated that by reduc¬ 
ing the severity of attack by the use of an 3^-in. vent 
and single-base powder, the microstructures of the 
bore surface layers became much more like those 
observed in eroded guns described in Section 12.1.2. 
The results of these tests were recorded 75 and the 
results of all vent tests were correlated with erosion 
rates in medium caliber guns. Under the less severe 
conditions pure cobalt and high-cobalt alloys includ¬ 
ing stellite-type alloys are superior in erosion resis¬ 
tance to gun steels while nickel alloys show inter¬ 
granular attack. 

The results of vent-plug tests have only a very 
limited applicability in the selection of materials to 
improve the performance of caliber .50 machine gun 
barrels and other automatic small arms barrels. The 
reasons for this are discussed and a new testing device 
is described in Section 16.3.6. Vent-plug tests have 
been somewhat more successful in the search for 
materials for the exhaust throats of recoilless 


Examination of Blocks After Firing in 
Explosion Vessels 

A modified erosion vent-plug apparatus, described 
in Section 11.2.4, was used to study changes produced 
in the surface of chromium and molybdenum as well 
as gun steels by the action of hot powder gases 
streaming past the flat surface of a block of the ma¬ 
terial to be tested, which had been set into a recess in 
the wall of the vessel. 31 Pressure was controlled by 
the size of a vent plug from which the gases issued 
after flowing past the D-shaped vents. Flat surfaces 
made possible ready microscopic and x-ray examina¬ 
tion of the blocks between successive firings. The 
results on gun steel are discussed in Section 12.2.2. 
The x-ray studies 137 of a chromium block after firing 
with double-base powder indicated a slight attack by 
powder gases with the formation of the hitherto un¬ 
known compound Cr 3 0 4 . In the case of molybdenum 
there was only a slight attack both with single-base 
and with double-base powders but the resultant prod¬ 
ucts could not be identified. 

Tests for Resistance to Surface Cracking 

A third type of explosion vessel (Section 11.2.4), 


similar in general construction to the two types just 
described, was used to study the types of surface 
crack patterns formed under simulated gun condi¬ 
tions when hot powder gases flowed across flat sur¬ 
faces of blocks of polished gun steel and of a number 
of other metals and alloys. 51 The specimen was ex¬ 
amined with a microscope to ascertain the effects 
produced. Surface cracks formed in pure iron and all 
steels, more rapidly in high-alloy steels than in gun 
steels. Intercrystalline cracks formed in pure nickel. 
No surface cracking occurred with molybdenum, 
chromium, tantalum or copper under the same con¬ 
ditions of test. 

In an attempt to separate as completely as possible 
the purely thermal effects of erosion from the chem¬ 
ical effects, an electron bombardment apparatus (de¬ 
scribed in Section 11.3.2) was designed whereby the 
surface of a small sample of gun steel or other mate¬ 
rial could be subjected repeatedly to a very high tem¬ 
perature of extremely short duration. The conclusion 
was reached that thermal action alone does not cause 
surface cracking, as discussed in Section 13.5.3. This 
apparatus has been described and the results of tests 
on gun steel and two other special steels have been 
recorded. 104 After bombardment a special steel, which 
was called “TEW alloy” (Ni 30, Cr 20, Mo 4, W 4, 
Ta 2, C 0.1, balance Fe) showed an expansion of the 
lattice in one direction. This method of study of 
erosion resistance might have value if the degree of 
control could be increased. 

16.3.2 Firing of Metal Particles 

into an Evacuated Tube 

Another method to ascertain the resistance of 
metals and alloys to chemical attack by hot powder 
gases was devised. It consisted in mixing fine particles 
(filings) of the sample to be tested with the propellant 
charge in a caliber .30 cartridge case, after which the 
cartridge was fired into an evacuated glass tube, the 
bullet emerging through a port in the forward end of 
the tube. The advantages of this technique were 
threefold. Not only did it determine erosion resist¬ 
ance but it was possible to identify the products of 
such attack and to detect grain growth or inversions 
that might be the results of exposure to hot powder 
gases, and thus affect the rate of erosion. The appa¬ 
ratus is described in Section 11.2.6. The residue de¬ 
posited on the walls of the tube was amenable to 
removal and examination by means of x-ray powder 
photographs which were compared with similar pho- 


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SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 


tographs of the original metal particles and other 
known standards. 

This technique was applied to a large number of 
substances. 79 Excellent resistance to structural and 
chemical attack by hot powder gases was shown by 
the pure metals cobalt, tungsten, and tantalum and 
by the following alloys: an iron-silicon alloy contain¬ 
ing 4.85% Si, Stellite No. 21, high-nickel alloys with 
silicon or aluminum or chromium, and Monel. Most 
nickel-iron alloys and the nickel steels, however, 
showed a slight increase in lattice size which may in¬ 
dicate some reaction with the hot gases. 

Only a very slight chemical attack was shoAvn by 
chromium, molybdenum, a chromium-tungsten alloy, 
K42B, Illium R, Nichrome V and nickel 70-tantalum 
30. Gun steels, other steels, pure iron, high-iron al¬ 
loys, and pure columbium were attacked by the hot 
powder gases. Powdered ferrosilicon was resistant to 
chemical attack by single-base powder and was found 
to retard oxidation of gun steel during firings with 
double-base powder, as mentioned in Section 15.6.4. 
Unfortunately all iron-silicon alloys are too brittle to 
be serviceable as liners or coatings in gun tubes. 

16-3,3 Cavitation Erosion Tests 

An investigation 38 of the relation of cavitation 
erosion to gun erosion was made by subjecting some 
31 samples of gun steels and other alloys, previously 
studied as erosion vent plugs (described in Section 
16.3.1), to magnetostriction oscillation in air-free 
water at 8,000 cycles per sec. The degree of correlation 
between the resistance to cavitation erosion and hard¬ 
ness of the samples was high. There seemed to be no 
correlation between the resistance to cavitation ero¬ 
sion and erosion caused by hot powder gases as de¬ 
termined in vent-plug tests. Further discussion of 
these tests is given in Section 11.3.3. 

16-3-4 Battelle Laboratory Gun 

During the years 1941 to 1943, the Battelle Me¬ 
morial Institute, under Army contract through Water- 
town Arsenal, undertook a survey of the resistance of 
metals and alloys to powder-gas erosion. An appa¬ 
ratus for smooth-bore testing, called the Battelle lab¬ 
oratory gun, 245,552 was developed. 

In the apparatus used in early tests, the material 
to be tested, in the form of a cylindrical tube, 1 in. 
long and 0.75 in. in inside diameter with a wall 
thickness of 0.27 in., was in contact with the gases 


from a propellant for a 37-mm gun for 0.0036 sec at a 
pressure of about 11,500 psi. A projectile was used in 
this case with the result that, in effect, the gases 
streamed through a vent of annular cross section. 
The part of the projectile that passed through the 
test specimen was a cylinder ground to have a dia¬ 
metral clearance of 0.005 in. It was screwed to a cast 
iron weight that slid in a frame and was subjected to 
a braking action after the projectile had traveled out 
of the tube. The erosion from the mechanical action 
of the projectile was insignificant. Therefore a pro¬ 
jectile was not used in a later form of the apparatus. 
Also, instead of using an annular orifice, a rectangu¬ 
lar cross section for the passage of the gases was 
achieved by mounting two flat surfaces of test speci¬ 
mens opposite each other at an optimum distance for 
the desired testing conditions. In this way a control 
sample could be fired with the unknown under identi¬ 
cal conditions. 

When the severity level of the various tests is taken 
into account, the results obtained with this apparatus 
were in close agreement with the results obtained 
with erosion vent plugs described in Section 16.3.1. 
Since the vent-plug tests are simpler, more rapid and 
cheaper to make, and since the Battelle gun evaluated 
mechanical factors only to a slight extent, Division 
1, NDRC, did not duplicate it. Instead, attention 
was concentrated on testing methods (described in 
the next four sections) which also measured the me¬ 
chanical suitability of a material as an erosion resist¬ 
ant liner or lining in a rifled gun bore. 

16-3-5 Caliber .50 Erosion Testing Gun 

The caliber .50 erosion-testing gun, 122 a special 
hypervelocity gun developed at the Franklin Insti¬ 
tute for an accelerated erosion test, has been de¬ 
scribed in Section 11.2.1. 

One of its most important uses was to determine 
the erosion resistance of a large number of metals and 
alloys under these hypervelocity conditions. A special 
form of the gun was developed to permit insertion of 
a short (8-in.) breech liner at the origin of rifling. 
Special materials were tested usually as rifled liners 
or as a plating or coating on the bore of a rifled gun- 
steel liner. A few full-length monobloc barrels with 
chromium electroplate applied to the whole bore sur¬ 
face were tested with regular or special projectiles. In 
most cases, a metallographic examination of the bore 
surface was made, at Harvard University, on the 
liner or barrel after firings. 


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TEST METHODS FOR EVALUATING PROSPECTIVE MATERIALS 


337 


These firing tests 76 - 77 aided in the evaluation of the 
materials discussed in Section 16.4, where references 
are given to other chapters for details about some of 
the more important ones. Of the materials tested as 
liners only molybdenum (particularly when hardened 
by a small amount of alloying with cobalt and by 
intensive mechanical working), tantalum, and chro¬ 
mium-base alloys gave promise of being sufficiently 
erosion resistant to withstand hypervelocity condi¬ 
tions. The caliber .50 erosion-testing gun was partic¬ 
ularly useful in the development of molybdenum 
liners. h Materials which gave similar promise in the 
form of coatings were pure chromium electroplates 
and duplex electroplates with a chromium bore sur¬ 
face and an undercoat of either pure cobalt or a 
cobalt-tungsten alloy (Section 20.2). 

Steels, Monel, and high-nickel alloys showed poor 
erosion resistance. The bore surface of liners of Stel¬ 
lite No. 21 and Stellite No. 22 melted when double¬ 
base powder was used, but showed excellent resist¬ 
ance to both melting and chemical attack with single¬ 
base powder. Pure cobalt and cobalt-tungsten alloys 
applied as electroplates melted with double-base pow¬ 
der and, in some cases, the latter were not sufficiently 
adherent, although both showed promise as under¬ 
coats under chromium electroplate. Copper plate 
melted; nickel-tungsten alloy electroplates, and Parco- 
lubrite coatings were not erosion resistant; duplex 
plates of nickel or copper or both under chromium 
and pyrolytic coatings of molybdenum on Stellite 
No. 21 or gun steel were mechanically unsatisfactory 
because of poor adherence or severe deformation. 

16 3 6 Short Rifled Caliber .30 Liners 

Because of the short duration of contact between 
powder gases and the bore surface, the high overall 
barrel temperatures during long bursts, (Section 
5.4.2) and the high rate of fire with the resultant 
severe mechanical effects of projectile impact on the 
bore surface, the results of erosion vent-plug tests 
(Section 16.3.1) have only very limited applicability 
in the selection of materials to improve the perform¬ 
ance of caliber .50 machine gun barrels and other 
automatic small arms barrels. Since the vent plug is 
not rifled and no projectile is used, the influence of 
various physical and mechanical properties, such as 

h Under a contract with the Bureau of Ordnance, Navy 
Department, additional tests in this weapon were carried out 
by the Franklin Institute in 1946 as part of the program for 
the further development of molybdenum liners (Section 18.1). 


ductility and hardness, on the durability of a ma¬ 
terial as a liner in any gun is not evaluated. 

To meet these objections a new testing device was 
developed. 81 In the first model a short rifled liner of 
the material to be tested served as a barrel and was 
fitted into the regular erosion vent-plug explosion 
chamber. Solid copper slugs were fired through the 
barrel and a rapid erosion rate was obtained by using 
a large powder charge. Results on the first trials gave 
excellent results which could be correlated with ero¬ 
sion in a caliber .50 machine gun. 

Since firing tests with this arrangement were slow, 
a new caliber .30 erosion-testing gun, designated 
Erosion Gun A, was designed. This fired at a much 
more rapid rate and employed readily available stan¬ 
dard caliber .30 AP bullets which were fired through 
a 6-in. long rifled barrel of the material under test. 
The barrel was mounted in a 37-mm breech mecha¬ 
nism, chambered for caliber .50 cartridge cases hand- 
loaded with double-base powder (13% nitroglycer¬ 
ine). Pressures were measured with a piezoelectric 
gauge threaded into the chamber. 

Erosion Gun A was used to test the relative per¬ 
formance of various electroplated coatings particu¬ 
larly chromium plates of various types and thick¬ 
nesses (Section 16.4.2), and was also used to test a 
liner bored from a rod of pure swaged cobalt 
(Section 16.4.7). 

16,3,7 Caliber .30 Machine Gun Liners 
or Barrels 

The first tests of possible erosion resistant materi¬ 
als in an actual gun under the erosion program of 
Division 1, NDRC were made by inserting short 
liners (some smooth bore and others rifled) of the 
special materials in the breech end of a caliber .30 
machine gun which was then fired with caliber .30 
bullets and double-base powder. The barrels were 
bore-gauged during the firing test and later subjected 
to metallographic examination. A high-alloy steel, 
pure nickel (Section 16.4.9) and steel liners with dif¬ 
fused chromium (Section 16.4.2) and with diffused 
tantalum (Section 16.4.6) bore surfaces were tested. 78 
All except the last showed a poorer resistance to 
erosion than gun steel under comparable conditions. 

Two high-alloy steels (Silchrome XCR and XB 
valve steels) that showed better erosion resistance 
and better hot hardness than gun steels, were selected 
for trial as caliber .30 aircraft machine gun barrels, 
as described in Section 16.4.11. 


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338 


SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 


16 3 8 Caliber .50 Heavy or Aircraft 
Machine Gun Liners or Barrels 

The test of materials as short breech liners in cal¬ 
iber .50 machine gun barrels or as electroplates or 
other bore coatings on short steel breech liners was 
found to be a very useful, convenient, and rapid 
method of determining the ability of a material to 
withstand the combined effect of high temperature, 
attack by powder gases, and mechanical stresses at 
the bore surface during the firing of this gun. 

A method of liner insertion for the testing of ma¬ 
terials was devised at the Geophysical Laboratory. 1 
Two caliber .50 heavy machine gun barrels were 
modified to make one barrel in two sections, a breech 
section containing the liner and a muzzle section, 
joined together with a threaded joint as described in 
Section 11.2.2. The caliber .50 Browning machine 
gun was selected because this gun could be fired read¬ 
ily and cheaply in a simple, easily constructed firing 
range under the control of the investigator. The 
heavy (28-lb, 45-in.) barrel was selected instead of 
the lighter (10-lb, 36-in.) aircraft barrel for trials of 
liner insertion because of the heavy barrel walls at 
the breech end of this barrel. 

Later the Crane Company found that the design it 
had developed for the stellite liner for caliber .50 air¬ 
craft machine gun barrels (Chapter 22) was satisfac¬ 
tory for testing other materials. The large number of 
firing results obtained at the Geophysical Laboratory 
and at Crane Company in caliber .50 machine gun 
barrels with liners or with plated bores or with com¬ 
binations thereof are given in two NDRC reports 80 81 
and the results are recorded in this report under the 
individual materials tested later, in this chapter and 
in Chapters 17 to 24, inclusive. 

Every gun, or at least every class of guns, is to 
some extent a separate problem with respect to req¬ 
uisite characteristics of its bore surface if erosion is 
to be minimized and performance enhanced. What 
would be a great improvement in a caliber .50 ma¬ 
chine gun barrel might be no improvement or might 


* The program of liner testing in the caliber .50 heavy ma¬ 
chine gun barrel undertaken at the Geophysical Laboratory 
combined three purposes: (1) to find out what happened in a 
barrel during firing (the mechanism of erosion and dimensional 
changes in the bore); (2) to obtain quickly some information 
on the relative erosion resistance and reaction to firing of 
various materials; and (3) to obtain data which might serve as 
a basis for predicting how erosion in larger guns might be 
minimized. (In connection with the latter purpose see the 
Introduction of Section 11.2.2.) 


be impractical to apply to a large antiaircraft gun or 
to an 8-in. Naval gun. Nevertheless, fundamental 
data can be obtained and much of it can be applied 
to the same size gun fired under any or all of the 
following conditions: more severe schedules, hyper¬ 
velocity conditions, higher cyclic rates of fire. Such 
information can also be applied to larger guns, par¬ 
ticularly if the significant differences between types 
of guns are studied and taken into consideration. 

Experiments performed in the development of very 
superior machine gun barrels showed that the best 
evaluation of the relative merits of various liner ma¬ 
terials could not be obtained if failure occurred by 
deterioration of the steel barrel into which the liner 
for test had been inserted. A satisfactory test of the 
liner material was only obtained when the liner failed 
before the barrel or when the liner and barrel “wore 
out” simultaneously. The roles of the liner, the bar¬ 
rel, and of an erosion-resistant electroplate on the 
steel bore ahead of the liner in the overall perform¬ 
ance of machine gun barrels are described in Chap¬ 
ter 24. The flexible CGL firing schedule which has 
been found so useful for the evaluation of materials 
and combinations and their performance is described 
in Section 24.1.3. 

16 3 9 Measurements of Thermal and 
Mechanical Properties 

In order to estimate whether there is any possibil¬ 
ity or probability that a given chemically resistant 
material might meet the thermal and mechanical 
requirements of a bore surface in a specific weapon, 
or whether the properties of such a material might be 
altered by metallurgical treatment (for instance, by 
alloying, mechanical working, or heat-treatment) to 
enable it to meet some requirement in which it may 
be deficient, a knowledge of certain thermal proper¬ 
ties and of the mechanical properties of the material 
at room temperature and at elevated temperatures is 
always desirable and usually necessary. Unfortun¬ 
ately, the part which the absence of or deficiency in a 
certain property may play in determining behavior 
and performance of the material in a gun is obscured 
by the interplay of factors. However, valuable pre¬ 
liminary information can be gained from tests of 
physical and mechanical properties and the choice of 
materials for firing tests can be narrowed down to a 
few promising substances. 

Fortunately, many of the thermal and mechanical 
properties of those pure metals and alloys, which 


CONFIDENTIAL 




TEST METHODS FOR EVALUATING PROSPECTIVE MATERIALS 


339 


were found to be resistant to chemical attack by hot 
powder gases, were known. However, many measure¬ 
ments on the most promising of these materials were 
made by contractors of Division 1, NDRC. These 
results are summarized in the other chapters of Part 
V where these specific materials are discussed. 

Meltability 

Whether or not a particular bore surface material 
reaches its melting point or softening range depends 
on the heat input under the particular firing condi¬ 
tions (Section 5.2), the rate of heat conduction away 
from the bore surface, and the specific heat of the 
bore-surface material. The relations of these factors 
have already been discussed in Section 5.4.4. The 
melting points of all the pure metals and the fusion 
ranges of most of the alloys were known as were their 
thermal conductivities and specific heats, and only 
a few additional measurements* were necessary. 

Thermal Expansion 

The ease or difficulty of satisfactory insertion of a 
liner of a special material or the adherence of an 
electroplate or other bore coating depends in part on 
the relative coefficients of thermal expansion of gun 
steel and the particular special material as well as the 
modulus of elasticity and other mechanical proper¬ 
ties. For example, major insertion difficulties must be 
surmounted in the satisfactory insertion of a molyb¬ 
denum liner, as discussed in Section 26.5.2. Because 
the coefficient of thermal expansion of molybdenum 
is only about one-half that of gun steel, a shrunk-in 
liner is no longer in heavy compression when the gun 
tube becomes hot and because of the high modulus 
of elasticity of molybdenum it takes a high propor¬ 
tion of the stress. 

The coefficient of thermal expansion of a material 
for use as the whole or a part of the bore surface of a 
gun tube should be low so that at the elevated tem¬ 
peratures reached during firing the rotating projec¬ 
tile does not lose engagement with the rifling and fail 
to attain adequate spin. Poor performance was ob¬ 
tained with special caliber .50 aircraft machine gun 
barrels of a stainless steel barrel material in which 
short breech liners of Stellite No. 21 had been in- 


j For data on the fusion range and specific heat of Stellite 
No. 21 and on the fusion range and thermal conductivity of 
chromium-base alloys, see Sections 19.3.8 and 17.3.1, respec¬ 
tively. 


serted, partly as a result of the high coefficient of ex¬ 
pansion of such a steel which resulted in inadequate 
spin and unstable bullets at a much earlier stage than 
with similar modified gun-steel barrels. 

Data on the thermal expansion of most materials 
were already available. Data on a number of materi¬ 
als were obtained at Crane Company 80 using a re¬ 
cording dilatometer. Measurements were made on 
chromium-base alloys (Chapter 17) at the University 
of Michigan for the Climax Molybdenum Company. 87 

Thermal Transformation 

Some pure metals or alloys may exist in two or 
more different crystalline forms. Each such form is 
stable only over a certain range of temperatures (and 
pressures) and when the inversion temperature (or 
inversion range) is reached there is an abrupt poly¬ 
morphic change accompanied by a volume change. 
Any abrupt volume change affects erosion resistance 
and in addition the relative erosion resistance of the 
several crystalline forms of a given material may be 
quite different. In the case of a coating on steel such 
volume changes during the alpha-gamma transfor¬ 
mation in the steel or changes in the coating at some 
inversion in the coating material may cause cracking, 
poor adhesion, and rapid deterioration of the bore 
surface. Thermal transformations in gun steels have 
already been discussed in Section 13.2.3. The inver¬ 
sion temperatures of cubic and hexagonal cobalt and 
cobalt solid solutions Avere determined and are dis¬ 
cussed in Section 19.3.2. Measurements 80 were made 
on many materials, using a recording dilatometer, in 
order to detect inversions and other phase transfor¬ 
mations that may occur in the heating and cooling 
cycle, as well as to measure the thermal expansion. 

Change in Properties on Annealing 
or Heat-Treatment 

Permanent changes in the mechanical properties of 
a bore surface material may result from exposure of 
such a material to the thermal cycle in a gun bore 
during firing. Hard chromium electroplate (Section 
20.2.2), for example, has a Vickers hardness number 
[VHN] of about 800 as deposited, and on annealing 
at elevated temperatures this hardness gradually 
drops to around 145 VHN. A similar decrease in the 
hardness of chromium plate occurs in plated guns as 
a result of annealing during the firing. In stellite and 
other alloys (see Section 19.3.5) a permanent change 


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340 


SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 


in hardness (precipitation- or age-hardening) may 
occur during firing, as well as the surface hardening 
(work-hardening) which may occur as a result of the 
mechanical working of the bore surface by the projec¬ 
tile. In addition to such changes in hardness and other 
mechanical properties, grain growth of metals or al¬ 
loys may also occur as a result of heating and cooling 
during firing. 

Tensile Strength, Ductility, and Modulus 
of Elasticity 

The behavior of erosion resistant metals and alloys 
as bore-surface materials, particularly when applied 
as relatively thick sections of metal in the form of 
gun liners, is dependent upon the strength of these 
materials especially at elevated temperatures. Unless 
an erosion resistant material is sufficiently strong to 
resist the high pressures as well as the high tempera¬ 
tures without a large permanent expansion or brittle 
failure, and strong enough at elevated temperatures 
to resist the mechanical forces of engraving of the 
projectile without permanent displacement of metal 
by flowage (swaging effect), it will not be a satisfac¬ 
tory bore-surface material. If the ductility of the 
material is low it will be unable to withstand the 
stresses that occur during the pressure and tempera¬ 
ture cycles in the bore, and also engraving stresses, 
without brittle failure by surface cracking apd 
cracking in depth. Both the strength and modulus of 
elasticity of the material as compared with gun steel, 
as well as the relative coefficients of expansion, deter¬ 
mine, in part, the adhesion of an electroplate or other 
bore coating in a steel bore and the ease or difficulty 
of insertion of a liner in a steel gun tube. 

Attention should be called here to the importance 
of high-temperature strength in the gun barrel mate¬ 
rial as well as in the erosion resistant bore-surface 
material. For use with severe firing schedules at or¬ 
dinary velocities or for hypervelocity conditions, the 
poor high-temperature strengths of conventional gun 
steels limit the performance that might be obtained 
by use of an erosion resistant liner or bore coating. 
No such liner or bore coating removes heat from the 
gun barrel. As described in Section 24.5, caliber .50 
aircraft machine gun barrels made of special heat- 
resisting steels, which also contain a breech liner of 
Stellite No. 21 and whose bores are chromium-plated 
ahead of the liner, show greatly improved perform¬ 
ance over similar lined and plated barrels made of 
WD 4150 steel. 


The results of measurements of ultimate tensile 
strength, yield strength, breaking strength, elonga¬ 
tion, reduction of area, and modulus of elasticity are 
summarized in Chapters 17 to 21, where specific ero¬ 
sion resistant materials are discussed. 

Notch Impact Tests and High-Velocity 
Impact Tests** 

When it is fired, a gun tube is subjected to rapidly 
applied stresses (Section 7.1), both from sudden large 
changes in pressure and also from projectile impact 
on the bore surface. Therefore it is desirable to obtain 
some estimate of whether an erosion resistant liner 
material might have sufficient toughness to prevent 
cracking, fracture, and disintegration under these 
conditions. For this purpose standard Charpy tests 
with either keyhole or V notches were made on some 
materials and the results are summarized under the 
specific materials in other chapters of Part V. The 
notch effect is important in determining the life of a 
material and the nature of its failure. For example, 
caliber .50 machine gun liners of Stellite No. 6, which 
(unlike Stellite No. 21) has a low Charpy value, gave 
excellent performance during initial" severe bursts of 
fire but later, presumably because surface cracking 
(heat-checking) became severe enough to provide 
“notches,” brittle failure occurred by cracking and 
propagation of cracks completely through the liner 
wall. 

More attention must be paid to the dynamic phys¬ 
ical properties of barrel steels at elevated tempera¬ 
tures, as well as to these properties of special bore- 
surface materials, in order to utilize more fully the 
advantages gained by the use of such an erosion re¬ 
sistant bore-surface material in the form of a breech 
liner or a plated bore. Firing tests, described in Sec¬ 
tion 24.5, have shown that regular barrel steels de¬ 
teriorate before an erosion resistant liner or coating 
shows any severe wear or other failure. 

Preliminary investigations were made at the Crane 
Company (with the apparatus shown in Figure 2) of 
the properties of steels under high-velocity impact 
conditions, comparable to the stress propagation con¬ 
ditions obtained during firing. These tests indicated 

k At Watertown Arsenal 277 a constant-deformation machine 
has been developed as a testing device for progressive stress 
damage in 8-in. tubes. It measures qualitatively the mutual 
influence of ductility of the barrel steel, shape of the stress- 
raiser, stress, rate of straining and temperature on the tenden¬ 
cy of steel to fail by cracking when loaded repeatedly for a 
relatively small number of times. 


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TEST METHODS FOR EVALUATING PROSPECTIVE MATERIALS 


341 



LOADING LEVER CATCH 



LOADING LEVER 


DUMMY SPECIMEN HOLDER 


CLUTCH LEVER 




TUBE SWITCH 


TUBE SWITCH 
LEVER 


STRIKER 

TENSILE 


TRIGGER 

LEVER 


SOLENOID 


CATCH 


PHOTOELECTRIC 

CELL 


Figure 2. 


Apparatus for high-velocity impact test. (Photographs used by courtesy of the Crane Company.) 


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SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 


strength values for metallic materials several times 
as great as the values accepted for static-design pur¬ 
poses. The mechanism of metallic deformation under 
conditions of high velocity and elevated temperature 
is of an essentially different nature than slow defor¬ 
mation. 1 Specifications for gun steels and for erosion 
resistant materials for liners should, therefore, be 
based upon the dynamic elevated temperature prop¬ 
erties, rather than upon the conventional criteria of 
chemical composition and static tensile properties at 
room temperature. Before such specifications can be 
prepared it must be ascertained which dynamic ele¬ 
vated temperature properties are pertinent to gun 
barrel performance and suitable apparatus for deter¬ 
mining such properties must be adequately devel¬ 
oped. It is recommended that such studies be pursued. 


are summarized for specific materials in other chap¬ 
ters of Part V. In a previous subsection on “Change 
in Properties on Annealing or Heat Treatment,” atten¬ 
tion has been called to the effects of the thermal cycle 
in a gun on the hardness of certain erosion resistant 
metals or alloys. 

A bore-surface material should have good resist¬ 
ance to abrasion and wear at high temperatures. 
Such resistance is presumably another corollary of 
high hot-hardness. No measurements of coefficient of 
friction or wear resistance were made. Measurements 
of these properties might have been helpful in select¬ 
ing bore-surface materials. The first problem is that 
of apparatus, a satisfactory form of which has not yet 
been devised. 


Hardness, Hot-Hardness, and Wear Resistance 

Firing tests on liners of pure molybdenum in the 
caliber .50 machine gun barrel (see Section 18.6.1) 
emphasized the importance of hot-hardness in an 
erosion-resistant liner material. Even though this 
soft metal was resistant to chemical attack by pow¬ 
der gases and to melting, early failure occurred by 
deformation of the rifling and its complete oblitera¬ 
tion by swaging impact of the bullets. Subsequent 
tests on barrels with chromium-plated bores (Section 
20.2.3) showed that the erosion resistance of chro¬ 
mium could not be fully utilized because the gun- 
steel lands beneath the chromium failed by deforma¬ 
tion and flowage unless they had been hardened by 
nitriding or induction-hardening. The rate of loss of 
hardness (and strength) of the steel bore beneath an 
erosion-resistant electroplate or other thin bore coat¬ 
ing plays an important part in determining the use¬ 
fulness and permanence of such a bore protection. 

The resistance of a material to deformation by the 
projectile is related to hot-hardness and hot-strength 
but is not determined by these alone but also by high- 
velocity impact resistance, mentioned in the preced¬ 
ing subsection. Hot-hardness data for most materials 
were not available and a large number of measure¬ 
ments were made at the Research Laboratory of the 
Climax Molybdenum Company for Division 1, 
NDRC contractors. Most of these results have been 
presented in several NDRC reports 59 109 110 and a few 

1 The influence of impact velocity on the properties of some 
metals and alloys at room temperature has been studied under 
the auspices of the War Metallurgy Committee (Division 18, 
NDRC). 558 


Hydrostatic Tube Testing 

A very useful method for determining the me¬ 
chanical behavior of erosion resistant materials was to 
test tubes with hydrostatic pressure. 81 By means of 
the simple apparatus shown in Figure 3 the tensile 
properties and other properties of a liner material in 



Figure 3. Hydrostatic tube-testing device. (This 
drawing was obtained by courtesy of the Geophysical 
Laboratory.) 


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EVALUATION OF LINER AND COATING MATERIALS 


343 


the shape of a tube could be determined. By electro¬ 
plating the inside of steel tubes, subjecting them to 
hydrostatic pressure of different amounts, and exam¬ 
ining them at different stages in the test, the inner 
fiber stress at which the plate cracked could be found. 

This method of testing was also used to detect flaws 
or lack of strength in individual liners before insertion 
in the gun barrel for test. A rifled liner was subjected 
to successively increased hydraulic pressures up to 
the maximum pressure of the gun barrel. If the liner 
passed this test but failed in the gun, such failure in 
all probability was not caused by a flaw in the par¬ 
ticular sample. This preliminary nondestructive test 
that insured the soundness of the liners fired was very 
helpful in the early stages of the development of 
molybdenum liners, before the fabrication methods 
described in Chapter 18 had made it possible to pro¬ 
duce samples without flaws. 

Adherence Test for Platings or Coatings 

To determine the relative adherence of platings or 
coatings prepared in different ways on steel or other 
material, a special ring test 93 was developed in which 
a ring cut from a plated tube was severely deformed 
by compressing it longitudinally at room temperature. 

Bend and Ring-Compression Tests 

In addition to the ring-compression test just de¬ 
scribed for determining the adherence of plates or 
coatings, similar ring-compression tests 96 were used 
to evaluate the ductility and strength of various 
samples of molybdenum fabricated by different meth¬ 
ods. During the development of improved fabrica¬ 
tion methods for molybdenum described in Section 
18.3, it was necessary to have a simpler test so that 
the strength and ductility of the metal could be 
determined quickly on a large number of specimens. 
A simple bend test 95 on a modified tensile and guided 
bend test machine was developed for this routine 
testing. 

Flaw Detection 

In addition to the use of the hydrostatic tube test¬ 
ing method (already described in this section) for the 
detection of flaws in a metal or alloy, several other 
methods were employed. Radiographic inspection 80 
of castings of Stellite No. 21 was used to detect blow¬ 
holes, porosity, and refractory inclusions. Cracks in 


molybdenum 95 were detected either by treating the 
pieces with aqua regia which caused the cracks to 
appear as black lines or by the use of the “zyglo 1 ’ 
method, where the piece is soaked in thin oil, sprinkled 
with a fluorescent powder, and then examined under 
ultraviolet light. Magnetic, electronic, and sonic de¬ 
vices for the detection of flaws in metals and alloys 
were considered but were not actually used in the 
erosion-resistant liner program. 

Machineability 

The machineability of a liner or gun barrel material 
is an important practical matter. Both tungsten and 
molybdenum were considered for development as 
possible liner materials for hypervelocity guns. One 
reason for rejection of tungsten in favor of molyb¬ 
denum was the lack of machineability of the former. 
Some alloys are considered difficult to machine be¬ 
cause no one has the “know how.” By the use of the 
proper cutting tools, feeds and speeds, at Crane Com¬ 
pany 119 it was easy and practical to bore and rifle 
liners of Stellite No. 21 and even liners of the much 
more refractory Stellite No. 6. Similarly, at Climax 
Molybdenum Company, satisfactory practical meth¬ 
ods of boring and rifling liners of chromium-base 
alloys 87 were developed. The ease with which molyb¬ 
denum may be machined was improved by studies 96 
made at Westinghouse Research Laboratories. 

There seems to be little doubt that, if necessary, 
grinding methods could be successfully developed for 
handling nonmachineable metals or alloys. Such 
methods can only be used for rifling gun bores of medi¬ 
um or large caliber. For small bores, such as caliber 
.30, .50, or .60 (or even 20 mm) grinding methods are 
probably impossible or at least impractical. 

16 4 EVALUATION OF LINER AND COATING 
MATERIALS 

16,41 Introduction 

By means of the test methods just described in 
Section 16.3, the possibilities of various metals and 
alloys as bore surface materials have been evaluated. 
No substance was found to possess the perfection 
required in an ideal erosion resistant material as 
described in Section 16.2. Several metals or alloys 
were found which, in addition to being resistant to 
chemical attack by hot powder gases, also met the 
thermal and mechanical requirements of a bore sur- 


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344 


SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 


face for a specific weapon under severe firing condi¬ 
tions or could be sufficiently changed by metallurgy 
(alloying, mechanical working, heat-treatment, etc.) 
to enable them to meet some requirement in which 
they were somewhat deficient. Before making an 
appraisal of possible bore-surface materials, a short 
analysis of the importance of the several factors con¬ 
tributing to failure of such materials will be given. A 
firing test 49 of a short breech liner of pure unhardened 
molybdenum in a caliber .50 heavy machine gun bar¬ 
rel showed that, even though there was little or no 
powder gas erosion (thermal and chemical attack), 
the lands at the origin of rifling and for some distance 
beyond were deformed and flattened, with the result 
that there was a very rapid loss in both velocity and 
accuracy. A similar result was obtained in regular 
steel barrels for this same gun, whose unhardened 
steel bores had been protected from chemical and 
thermal attack by an erosion-resistant plating (chro¬ 
mium electroplate). The steel beneath the plate de¬ 
formed and rifling was obliterated, as described in 
Section 10.5.3. These results emphasized the impor¬ 
tance of hot-hardness, resistance to permanent de¬ 
formation, and wear resistance in a machine gun 
barrel, even at ordinary velocity. 

The hypothesis was advanced that, since the time 
that powder gases are in contact with the bore sur¬ 
face is very short in the caliber .50 machine gun, 
possibly hot-hardness and wear resistance rather 
than resistance to powder gas erosion might play the 
most important part in determining performance. To 
test this hypothesis a few materials of known hot¬ 
hardness but different composition were selected for 
test. Liners of high speed steel, two hot die steels, and 
Stellite No. 6 were prepared and tested. 80 The first 
three materials all showed poor performance, owing 
to thermal and chemical attack by powder gases and 
two of them cracked besides. The liner of Stellite No. 
6, on the other hand was resistant to thermochemical 
attack and to deformation of the rifling. However, 
ultimate failure occurred by cracking. These results 
showed that resistance to thermochemical attack was 
a prime requisite for a bore surface material, even in 
the caliber .50 barrel at ordinary velocity. It turned 
out that in the stellite selected for this test, hot¬ 
hardness had been overemphasized at the expense of 
ductility, for later tests of Stellite No. 21 showed that 
it has adequate hot-hardness and wear resistance 
combined with excellent ductility. The outstanding 
success with this material as a machine gun liner, will 
be outlined later in Section 16.4.8. 


The comparative performance of Stellite No. 21 
and molybdenum in caliber .50 machine gun barrels 
and of Stellite No. 21 and molybdenum (both un¬ 
hardened and hardened) under hypervelocity con¬ 
ditions in the caliber .50 erosion-testing gun showed 
that resistance to plastic deformation and wear are 
important qualities of an erosion resistant material 
but are not sufficient to meet all of the conditions. 
The deformation of the rifling in liners of pure un¬ 
hardened molybdenum in both of these guns showed 
that resistance to powder gas erosion (thermal and 
chemical attack) is also not a self-sufficient quality. 
The very promising performance 70 of a liner of hard¬ 
ened molybdenum (Section 18.6.2) in the caliber .50 
erosion-testing gun indicated that the combination of 
thermochemical resistance (no chemical action with 
or melting by hot powder gases) with high hardness 
at the working temperature of the gun is a sufficient 
qualification provided, however, that the liner ma¬ 
terial has sufficient ductility and impact resistance 
both hot and cold to prevent brittle failure and dis¬ 
integration. Preliminary firing results 87 with liners of 
chromium-base alloys (Section 17.4.3) indicated that 
even relatively brittle materials, which have the other 
necessary requisites of a bore-surface material, may 
be utilized if they are so inserted as to maintain them 
under compression and minimize the possibility of 
brittle impact failure. 

An appraisal of possible bore-surface materials will 
now be given. Additional details about the most 
promising ones are given in later chapters of this 
report. For further information about the others ref¬ 
erences are given to the reports submitted by the 
contractors of Division 1. 

16 4 2 Chromium Electroplates, Duplex Plates 
and Other Chromium Coatings 

Chromium electroplates show considerable promise 
as bore-surface materials both at conventional ve¬ 
locities with high rates of fire and under hypervelocity 
conditions. Pure chromium deposited electrolytically 
is a brittle metal and quickly develops cracks as a 
result of firing stresses and some recrystallization 
during firing. Hot powder gases penetrate these 
cracks and attack the underlying material (see Sec¬ 
tion 13.2.3) and possibly the chromium and finally 
there is failure by undercutting, as described in Sec¬ 
tion 20.2.1. As a result of the studies of Division 1, 
NDRC, combined with those undertaken by the 
Army Ordnance Department, the effectiveness of 


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EVALUATION OF LINER AND COATING MATERIALS 


345 


chromium plate in resisting erosion and thus improv¬ 
ing the life and performance level of automatic, 
rapid-fire guns has been demonstrated and the effec¬ 
tiveness of chromium plate for service under hyper¬ 
velocity conditions has been indicated. 

The possible advantages of duplex plates and stud¬ 
ies of alloy plates are discussed in Chapter 20. The 
development of improved machine gun barrels by the 
application of chromium electroplates in conjunction 
with bore hardening or with liners and other barrel 
modifications is described in Chapters 23 and 24. The 
commercial development of the chromium plating of 
small arms barrels is described in Chapter 25. The 
success obtained in improving plate utilization, partic¬ 
ularly under hypervelocity conditions, by mitigation 
of the mechanical effects of the engraving of the pro¬ 
jectile by the use of pre-engraved projectiles or by 
the use of the Fisa protector is described in Chapters 
31 and 32, respectively. 

In addition to the use of hardened steels under 
chromium electroplate, “hot-hard” liners of Z nickel 
and of Stellite No. 21 were tried as base material for 
chromium electroplate. The results were unsatisfac¬ 
tory because of poor adhesion of the plate on these 
surfaces. 

The deposition of chromium pyrolytically from the 
vapor of its carbonyl was investigated, 74 93 as de¬ 
scribed in Section 21.6. The method does not appear 
to be suitable for gun bores. The carbonyl 73 is difficult 
to prepare. The process of plating is complex and 
difficult and the metal, which contains some carbon, 
is hard and very brittle. 

A firing test was made in a caliber .30 machine 
gun 78 on a steel liner with a bore surface of diffused 
chromium prepared by heating it in an atmosphere 
of chromium chloride. 322 Serious erosion occurred and 
performance was inferior to gun steel. Attention is 
called to the point that a “diffused chromium” sur¬ 
face is not a surface of pure chromium but a graded 
series of iron-chromium alloys produced by a diffu¬ 
sion or “cementation” process. 

As is brought out in Section 20.5, investigation 
should be continued concerning the chromium elec¬ 
troplating process, types of pure chromium plate, 
alloy plates and duplex plates, and the properties of 
such plates, and the effects of gun tube and projectile 
design on their performance. Studies should be made 
of the performance in medium-caliber guns of these 
plates on hardened steel bores and on special full 
length or partial length thin bore liners or linings of 
“hot-hard” alloys. 


16 4 3 Chromium and Chromium-Base 
Alloys 

Erosion vent-plug tests, other laboratory tests, and 
examination of worn chromium-plated medium-cal¬ 
iber guns had emphasized the excellent resistance of 
chromium metal to melting and chemical attack by 
hot powder gases even under severe conditions. At¬ 
tempts were made to utilize this metal as a liner 
material. The outstanding deficiency of chromium is 
cold ductility. Hence a liner made from chromium 
electroplate (Section 17.2.2) cracked after having 
been fired only a few rounds. 49 

Pure chromium, when melted and cast in a vacuum, 
yielded ingots of soft but brittle metal. Attempts 83 
to impart ductility by control of purity, by mechan¬ 
ical working, and by additions of small to moderate 
amounts of alloying constituents, which are described 
in Chapter 17, were unsuccessful. Although the metal 
showed some hot malleability over a limited range of 
temperatures, it was always brittle when cold. 

However, a group of chromium-base alloys 87 (vac¬ 
uum melted alloys with approximately 60% chro¬ 
mium, 25 to 30% iron, 10 to 15% molybdenum with 
very low carbon) shows considerable promise. They 
have very good resistance to thermal and chemical 
attack under hypervelocity conditions and are strong 
and hard so that no deformation of the rifling occurs. 
Their one weakness is their low ductility. In spite of 
this deficiency, when liners were properly inserted in 
caliber .50 aircraft machine gun barrels, they showed 
outstanding performance and did not crack and dis¬ 
integrate. These alloys and tests are discussed in 
Chapter 17. 

Since both chromium metal and copper have ex¬ 
cellent resistance to chemical attack by hot powder 
gases, an attempt was made to improve the mechan¬ 
ical properties of chromium by preparing powder 
metallurgy compacts 83 of chromium powder impreg¬ 
nated with an optimum amount of copper with or 
without subsequent mechanical working by swaging. 
The best of these compacts showed low weight losses 
(considerably less than gun steels) in erosion vent- 
plug tests 49 but not as low as were shown by corre¬ 
sponding compacts made from molybdenum or tung¬ 
sten powders (Sections 16.4.4 and 16.4.5). Because of 
the melting out of copper and the low hardness of the 
compacts, they cannot be considered potentially use¬ 
ful as gun liner materials. Nickel and palladium were 
tried as impregnating materials for chromium powder 
but attacked the chromium and formed alloys. 


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846 


SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 


16 4 4 Molybdenum and its Alloys 

Pure molybdenum has outstanding resistance to 
both thermal and chemical attack by powder gases 
even under the most severe conditions. The melting 
point (2620 C) is so high as to remove the possibility 
of bore-surface melting. Firing tests 49 ’ 76 showed, how¬ 
ever, that as usually fabricated it was not strong 
enough, not hard enough, and not ductile enough to 
make a satisfactory liner. Fabrication methods had 
not been developed to permit preparing the desired 
shapes and sizes. Furthermore, owing to the large 



Figure 4. Helical two-stave liner of molybdenum 
containing 0.1 per cent cobalt fired 2,024 rounds with 
ball bullets, M2 and double-base powder in the caliber 
.50 erosion-testing gun. Cross section y 2 in. beyond 
origin of rifling. Etched with 10% KOH+10% K 3 Fe 
(CN)6j 100X. (This figure has appeared as Figure 11 in 
NDRC Report No. A-405.) 

difference in coefficients of thermal expansion and 
modulus of elasticity between molybdenum and gun 
steel, liners are difficult to insert. 

As described in Chapter 18, development 52 ’ 95 96 has 
shown that all three of the mechanical deficiencies 
can be overcome by alloying with small amounts 
(about 0.1 to 0.2%) of cobalt or certain other metals 
and by intensive mechanical working accompanied 
by proper heating and annealing schedules. Satis¬ 
factory fabrication and insertion methods have been 
developed. Suitably fabricated molybdenum alloy 
mo^t nearly approaches the properties of the ideal 
erosion resistant material described in Section 16.2. 
Figure 4 shows a cross section of a liner of molyb¬ 
denum containing 0.1 per cent cobalt, which had 


been fired 2,024 rounds. The bore surface of this same 
liner is shown in Figure 19 of Chapter 18. Sufficient 
progress in the development of hardened molybden¬ 
um was made under Division l’s auspices, so that the 
Navy Department continued to support the efforts 
of the Westinghouse Electric Company to make 
this material in a form suitable for large gun liners, 
following plans developed by Division 1, as described 
in Section 33.1.3. 

In addition to fabrication by powder metallurgy 
methods, studies 97 were made of arc melting, thermit 
melting, and vacuum melting of molybdenum and 
high-molybdenum alloys. These methods, which do 
not seem to offer any advantages over powder metal¬ 
lurgy, are also described in Section 18.7. Pyrolytic 
plating 93 of molybdenum from the vapor of its car¬ 
bonyl was investigated and, as described in Section 
21.3, does not appear to be suitable for gun bores. 

Aqueous electroplating 80 of molybdenum was in¬ 
vestigated, but the deposits were impure and poorly 
adherent. Molybdenum coatings on steel were also 
made by depositing the suboxide from aqueous solu¬ 
tions and then heat-treating in hydrogen. 95 These 
plates were brittle and the process was time-consum¬ 
ing and generally unsuitable. Sprayed coatings 90 on 
steel bores were prepared and shown by firing tests 76 
to be brittle and completely unsatisfactory. 

In an attempt to avoid fabrication difficulties and 
improve ductility, powder metallurgy compacts of 
molybdenum powder and copper (trade name Elko- 
nite) were prepared and tested. 49 They showed excel¬ 
lent resistance to powder gas erosion, nearly com¬ 
parable to that of pure molybdenum in tests as 34-in* 
erosion vent plugs. However, when one of the best 
molybdenum-copper Elkonites was tested as a short 
rifled tube (See Section 16.3.6) using copper slugs as 
projectiles, the tube split when tested without a steel 
backing and showed deformation when adequately 
supported to prevent splitting. These alloys lack the 
strength and hardness to resist swaging impact by 
the projectile and show little promise as gun liners. 

16 4 5 Tungsten and Its Alloys 

Tungsten has the highest melting point of any of 
the erosion resistant metals. Because its melting 
point is higher than the adiabatic flame temperatures 
of conventional propellants, no melting of a tungsten 
bore surface could occur. All laboratory tests showed 
that pure tungsten and all tungsten-molybdenum 
alloys have excellent resistance to chemical attack by 


CONFIDENTIAL 





EVALUATION OF LINER AND COATING MATERIALS 


347 


powder gases. In spite of these advantages, no at¬ 
tempt was made to develop tungsten as a liner ma¬ 
terial because it was short in supply, presented the 
same fabrication problems (only in a more serious 
degree) as molybdenum, had the same mechanical 
deficiencies and difficulties of liner insertion as molyb¬ 
denum, and besides tungsten (unlike molybdenum) 
is machineable only with very great difficulty. 

Pyrolytic plating 93 of tungsten and of tungsten- 
molybdenum alloys on steel gun bores was investi¬ 
gated (Section 21.4) and found to be unsuitable be¬ 
cause of the formation of brittle intermetallic com¬ 
pounds at the interface between steel and the plate 
during firing. Sprayed bore coatings 90 of tungsten 
were prepared. These were brittle and fragile and 
never subjected to firing tests. 

Like the corresponding powder metallurgy com¬ 
pacts of molybdenum or chromium powder and cop¬ 
per, the tungsten-copper Elkonites were found to be 
resistant to powder gas erosion but even when the 
copper content was optimum for mechanical proper¬ 
ties, they were unsuitable for gun liners because of 
lack of sufficient strength and hot-hardness to resist 
the swaging impact of the projectile. 

16 4 6 Tantalum and Diffused Tantalum 

Of the metals resistant to chemical attack by hot 
powder gases, only tungsten has a higher melting 
point than tantalum. However, pure tantalum is 
very soft—nearly as soft as pure copper—but, unlike 
copper, it is readily hardened by air, hydrogen, nitro¬ 
gen, or any gases containing oxygen, nitrogen, hydro¬ 
gen, or carbon. No hot-hardness data are available on 
hardened tantalum, and there is some doubt whether 
its hardness and wear and galling resistance are ade¬ 
quate for rapid-fire guns. Excellent promise for 
hypervelocity service in guns was confirmed by a 
firing test 76 on a rifled liner of air-hardened tantalum 
in the caliber .50 erosion-testing gun. The cross sec¬ 
tion in Figure 5 shows some cracking but no evidence 
of powder gas erosion. A liner for a caliber .50 ma¬ 
chine gun barrel was prepared but was badly torn 
during an attempt to rifle it and never reached the 
test stage. 

Tantalum is machineable with some difficulty using 
carbon tetrachloride as a lubricant but is easily galled 
or torn. In spite of its promise as a bore-surface 
material under hypervelocity conditions, no further 
development of tantalum as a liner material was 
made by Division 1 because it was so expensive, so 


short in supply, and so strategic for other war uses. 

An attempt was made to plate tantalum on steel 
from a fused salt bath of potassium fluoride, po¬ 
tassium tantalum fluoride, and tantalum oxide at 
about 800 C. The results were erratic and no sa¬ 
tisfactory plates were produced. 

In firing tests 78 in the caliber .30 machine gun, two 
rifled steel liners with diffused tantalum bore surfaces 
of different thicknesses (0.0005 in. and 0.001 in.) 
performed better than gun steel. The thinner coating 
was more adherent and effective. Attention is called 
to the nature of diffused tantalum surfaces: such 
surfaces are not pure tantalum but a graded series of 
tantalum-iron alloys. 



Figure 5. Rifled tantalum liner air hardened and fired 
150 rounds with ball bullets M2 and double-base powder 
in the caliber .50 erosion-testing gun. Longitudinal sec¬ 
tion of a groove about 3^ in. from the origin of rifling. 
The cracks appear wider than they really were, because 
they became broadened during the polishing of the sec¬ 
tion. Etched with aqueous solution of 25% H 2 F 2 and 
25% EbSCh; 200X. (This figure has appeared as Figure 
12 in NDRC Report No. A-405.) 

16 4 7 Cobalt and High-Cobalt Alloys 

Pure cobalt and high-cobalt alloys show excellent 
resistance to chemical attack by hot powder gases. 
Unfortunately, their relatively low melting points 
(1490 C for pure cobalt and lower fusion ranges for 
the alloys) limit their application under the most 
severe conditions where bore melting is an important 
factor in erosion. Pure cobalt has insufficient hot¬ 
hardness to resist deformation by swaging impact of 


CONFIDENTIAL 




348 


SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 


the projectile in rapid-fire guns but most of the 
high-cobalt alloys have adequate hot-hardness and 
excellent strength and ductility. 

Pure cobalt was tested 59 as a short rifled liner in the 
caliber .30 erosion gun “A” (see Section 16.3.6). Al¬ 
though it showed excellent resistance to chemical 
attack and melting, deformation of the rifling occur¬ 
red. A liner of a cobalt alloy containing 7% tungsten 
tested 81 in a caliber .50 heavy machine gun barrel 
showed excellent performance, but is probably no 
better than Stellite No. 21 (see Section 16.4.8), which 
contains much less cobalt and can be fabricated more 
easily. A cobalt-base alloy (63 Co, 32 Fe, 4 Cr, 0.02 C) 
showed better erosion resistance than gun steels in a 
vent-plug test with 3^-in. vent and showed good hot¬ 
hardness up to 600 C, but its hardness deteriorated 
very rapidly above this temperature and no further 
tests of this alloy were made. The results of lab¬ 
oratory tests on iron-cobalt alloys are described in 
Section 16.4.11. 

A gun-steel liner electroplated with pure cobalt 
0.005-in. thick was tested 76 in the caliber .50 erosion¬ 
testing gun under hypervelocity conditions with 
double-base powder. Adherence was satisfactory but 
melting and deformation of the rifling occurred. 
Cobalt-tungsten alloy plates (see Section 20.4), 
which can readily be hardened by aging, were de¬ 
posited on gun-steel liners and tested both in the 
caliber .50 erosion-testing gun under hypervelocity 
conditions and in the heavy and aircraft caliber .50 
machine gun barrels. Melting occurred under hyper¬ 
velocity conditions and difficulties with adhesion 
were encountered in all tests unless the plate had 
been bonded to steel by a high-temperature diffusion 
treatment. These plates require more development 
and show considerable promise as bore-surface 
materials in medium caliber guns at normal veloc¬ 
ities. The promise shown by pure cobalt plates 
and the alloy plates as undercoatings beneath chro¬ 
mium plate has already been pointed out in Section 
16.4.2. 

16 4 8 Stellite No. 21 and Other Stellites 

The stellites are cobalt-base alloys with 25 to 30% 
chromium, and with a moderate amount (usually 5%) 
of molybdenum or tungsten, and small or moderate 
amounts of one or both of the metals iron and nickel, 
and with carbon contents varying from 0.2 to slightly 
over 1%. Stellites, because of their resistance to 
chemical attack by hot powder gases, hardness, and 


strength at high temperatures, and excellent wear 
and abrasion resistance, have shown outstanding per¬ 
formance in applications as short breech liners in 
machine gun barrels under firing conditions so severe 
that unmodified steel barrels are unable to withstand 
such schedules. 

A large number of stellites have been tested as 
liners but, as shown in Chapter 19, none is as satis¬ 
factory as Stellite No. 21, which has the optimum 
combination of mechanical properties (including both 
hot-hardness and good ductility) for service as a liner 
material. The outstanding deficiency of the stellites 
is their low melting ranges (usually 1250 to 1300 C) 
which limits their application to guns in which bore 
melting is not an important factor in erosion. In the 
caliber .50 erosion-testing gun 76 under hypervelocity 
conditions liners of Stellite No. 21 performed satis¬ 
factorily with single-base (IMR) powder but the 
bores melted when double-base powder was used, as 
illustrated in the two parts of Figure 6. 

Attempts were made to utilize the good ductility 
and hot-hardness of Stellite No. 21, under hyperve¬ 
locity conditions or severe firing conditions in ma¬ 
chine guns at an increased velocity obtained by the 
use of double-base powder, by protecting a stellite 
bore with a high-melting erosion resistant plate. 
These attempts were unsuccessful. Chromium depos¬ 
ited electrolytically from aqueous solutions had poor 
adherence and molybdenum applied by pyrolytic 
plating 93 was unsatisfactory because of the formation 
of a brittle intermetallic compound at the interface 
between plate and base material during firing, as 
described in Section 21.3. Further efforts should be 
made to deposit chromium by electroplating and 
then improve adhesion by high temperature bond¬ 
ing. Nitriding had no effect on the performance of 
liners of Stellite No. 21. 

The properties of the stellites (particularly Stellite 
No. 21), the development of Stellite No. 21 as a liner 
material, studies of the utilization of this alloy (cast¬ 
ing methods, bonding to steel, and the effects of hard¬ 
ness, composition, heat-treatments, method of inser¬ 
tion, liner length and wall thickness on performance), 
tests of stellites under hypervelocity conditions, ap¬ 
plications of Stellite No. 21 to the 37-mm M3 cannon, 
and tests of other stellites in machine gun barrels are 
described in Chapter 19. Figure 7 shows a cross sec¬ 
tion of a Stellite No. 21 liner which had been fired 
almost 1,100 rounds in a caliber .50 machine gun. The 
successful insertion of liners of Stellite No. 21 and 
their application to machine gun barrels, and their 


CONFIDENTIAL 



EVALUATION OF LINER AND COATING MATERIALS 


349 



Figure 6. Investment cast liners of Stellite No. 21 
fired with ball bullets, M2 in the caliber .50 erosion¬ 
testing gun; (A) 509 rounds with I MR (single-base) 
powder; (B) 85 rounds with FNH-M2 (double-base) 
powder. Cross sections 2 in. from the origin of rifling. 
The edge of the land in (A) is cracked; whereas the cor¬ 
responding feature in (B) was melted away. Etched with 
10% KOH + 10% K 3 Fe(CN) 6 ; 150X. (This figure has 
appeared as Figures 13 and 14 in NDRC Report No. 
A-405.) 


adoption for Service use are described in Chapter 22. 
The combination of Stellite No. 21 liner with choked- 


muzzle chromium plating ahead of the liner and other 
barrel modifications (weight, contour and composi¬ 
tion of barrel steel) to effect even greater increases in 
performance is described later in Chapter 24. 



Figure 7. Cross section of a nine-inch investment cast 
Stellite No. 21 liner fired 1,098 rounds in a caliber .50 
aircraft machine gun barrel at Purdue University range. 
Electrolytic aqua regia etch; 100X. (This figure has 
appeared in NDRC Report No. A-416.) 

16.4.9 Nickel and Nickel Alloys 

Pure nickel has too low a melting point (1452 C), is 
too soft, and is not inert enough chemically for service 
as a bore-surface material under severe firing condi¬ 
tions. Even when hardened (and certain high-nickel 
alloys show excellent hot-hardness) 59 the performance 
is still unsatisfactory. Firing tests on pure nickel and 
high-nickel alloys, as gun liners and as electroplates, 
and laboratory tests, showed that these (except 
nickel-chromium alloys with more than 10% chro¬ 
mium) were subject to severe intergranular attack by 
powder gases, illustrated in Figure 8, and gave a per¬ 
formance barely equal to or inferior to gun steel. 

One supposition 78 was that this intergranular at¬ 
tack was caused by the presence of small amounts of 
sulfur in the powder gases. It may have been related 
to a very small increase in the lattice dimensions of 
fine particles of nickel and some high-nickel alloys 
that were mixed with the propellant charge and fired 
into an evacuated tube (see Section 16.3.2). The reason 
for this increase was not investigated but it seems quite 
probable that it is caused by the solution of some 
carbon. 79 Similar experiments disclosed that high- 


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350 


SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 


nickel alloys containing silicon, aluminum, chromium, 
or copper (Monel metal) were unaffected by the hot 
powder gases. The high-nickel alloys with silicon or 
aluminum were not tested further and the results 
on the other alloys when tested as liners are de¬ 
scribed below. 

The nichromes with more than 10% of chromium 
were resistant to intergranular attack but were too 
soft to withstand deformation by the projectile and 
when hardened (for example the Inconels) they lose 
their erosion resistance. Hastelloy “C” (an alloy with 



Figure 8. Surface of a rifled nickel liner fired 500 
rounds with double-base powder in a caliber .30 barrel. 
100X. (This figure has been taken from a joint progress 
report on Contracts OEMsr-537 and OEMsr-608.) 

about 58% nickel, 17% molybdenum, 15% chro¬ 
mium, 4% tungsten and 6% iron) shows promise 80,136 
as a liner material in machine gun barrels or under 
conditions where bore melting is not an important 
factor in erosion. The first liner of this alloy tested in 
the caliber .50 aircraft machine gun barrel was hard¬ 
ened for maximum hardness and failed by cracking. 
The second liner was tested in the dead soft condition 
and failed by swaging of the rifling. Termination of 
the experimental program of Division 1, NDRC, after 
V-J day prevented further tests at intermediate hard¬ 
nesses, which should be completed. 

Tests of electroplated nickel in the caliber .50 
machine gun showed that its erosion resistance was 


about equal to that of gun steel, 49 which confirmed 
earlier British experience with medium-caliber guns. 
A number of nickel-tungsten alloy electroplates were 
tested in the caliber .50 erosion-testing gun 76 and in 
the caliber .50 machine gun. 59 All showed poor resist¬ 
ance to powder gas erosion and rapid removal in a 
few rounds. 

An alloy designated “Z” nickel, which contains an 
unspecified hardening agent, and zirconium-nickel 
(with about 0.5% zirconium) were tested as liners, 
both in the caliber .50 erosion-testing gun 76 and in the 
caliber .50 machine gun, 49 and showed an erosion rate 
greater than that of gun steel and a characteristic 
type of erosion (different from gun steel) which was 
caused by intergranular attack and detachment of 
grains of metal from the bore. A liner of Monel metal 
(67.5% nickel, 28.5% copper) showed the same type 
of attack and poor performance in the caliber .50 
erosion-testing gun. It was demonstrated 50 that the 
alloying elements in these liners just described were 
not responsible for the erosion by test in a caliber .50 
machine gun barrel of a liner of the purest nickel 
commercially available. In this test the erosion of 
nickel involved both melting and removal of grains 
of nickel following weakening along grain boundaries 
(intergranular attack). A previous test of a liner of 
pure nickel in the caliber .30 machine gun 78 showed 
performance inferior to gun steel. 

An even more conspicuous failure as a gun liner 
material was Colmonoy No. 5, a high-nickel alloy 
(78% nickel) containing some chromium, iron, silicon, 
and boron, which has been a competitor of stellites 
for cutting tools and valve seats for high temperature 
service. When tested 80 as a liner in the caliber .50 
heavy machine gun barrel it failed both by cracking 
and general melting of the bore surface, when fired 
only 45 rounds. Firing tests 80 in the caliber .50 air¬ 
craft machine gun barrel on liners of Hastelloy “A” 
and Hastelloy “B”, which are nickel-molybdenum- 
iron alloys (with no chromium present), showed that 
in addition to being too soft they were subject to 
chemical attack by hot powder gases. 

16410 Hardened 

Iron-Nickel-Cobalt-Chromium Alloys 

The remarkable performance of Stellite No. 21 as 
a liner material in machine gun barrels, which has 
already been outlined in Section 16.4.8 of this report, 
focused attention on the hot-hardness of a bore-sur¬ 
face material as an important requisite of such a 


CONFIDENTIAL 





EVALUATION OF LINER AND COATING MATERIALS 


351 


material. Because of the possibility of a critical short¬ 
age of cobalt in case the war should be of long dura¬ 
tion, an attempt was made to find a satisfactory 
substitute for Stellite No. 21 (which contains about 
64% cobalt) with a lower cobalt content or no cobalt. 
Because of these considerations, in addition to stud¬ 
ies of the effect on liner performance of replacing a 
part of the cobalt of Stellite No. 21 with one or both 
of the metals nickel and iron, a preliminary recon¬ 
naissance was made of the possibilities of cast or 
wrought iron-nickel-cobalt-chromium alloys with low 
or moderate carbon contents and any or all of the 
metals, molybdenum, tungsten, columbium, and tan¬ 
talum, as hardeners. 

Fortunately, about 77 heat-resisting alloys of this 
series were under investigation and development by 
the War Metallurgy Committee (Division 18, 
NDRC), the Navy Department and the National 
Advisory Committee for Aeronautics for use as blades 
in gas turbines, in turbosuperchargers and for other 
important high temperature uses. In order to expedite 
the preliminary study of the potentialities of these 
alloys as materials for gun liners, an arbitrary separa¬ 
tion into groups was made as follows on the basis of 
chemical composition : 

Group I. Stellite-type alloys: Cobalt-chromium 
alloys (cobalt-chromium ratio roughly 70 Co-30 Cr) 
with about 5% of either tungsten or molybdenum 
and varying carbon content from 0.1 to 0.5%. Part 
(up to about one-half) of the cobalt may be replaced 
by nickel. 

Group II. K 42-B-type alloys: Low-carbon al¬ 
loys, approximately Ni 40, Co 25, Cr 20 with Ti and 
A1 as hardening agents. 

Group III. Refractaloy-type alloys: Co-Ni-Cr al¬ 
loys with about 12% of body centered cubic metals 
(Mo, W, Cb, Ta) and varying carbon contents be¬ 
tween 0.05% and 0.5%. (For example: 30 Co, 20 Ni, 
20 Cr, 15 Fe, 8 Mo, 4 W, 0.1 C.) 

Group IV. N-155-type alloys: Co-Ni-Cr alloys 
with about 6% of body centered cubic metals (Mo, 
W, Cb, Ta) and varying carbon contents. (For ex¬ 
ample : 20 Co, 20 Ni, 20 Cr, 30 Fe, 3 Mo, 2 W, 1 Cb, 
0.3 C.) 

Group V. S-497-type alloys: Co-Ni-Cr alloys 
(lower Cr than Refractaloy or N-155-types) with 
about 12% of body centered cubic metals (Mo, W, 
Cb) and varying carbon contents between 0.1% and 
0.5% (For example: 20 Co, 20 Ni, 14 Cr, 30 Fe, 4 Mo, 
4 W, 4 Cb, 0.4 C.) 

Group VI. Hardened nickel-chromium alloys: 


Modified Inconels with hardening agents such as Be, 
Ti, Al, Mo, W. 

Group VII. Hardened stainless steels: Modified 
18-8, 25-12 or plain Cr stainless steels with hardening 
agents such as Ti, Mo, W, Cb, Mn. 

Some members of all except the last type show 
promise as possible gun-liner materials. Most hard¬ 
ened nickel-chromium alloys gave poor test results 
except Hastelloy “C” (see Section 16.4.9). Of the 
remainder, K 42-B-type alloys show the least prom¬ 
ise for gun liners. 

A reconnaissance series of these hot-hard alloys 
was prepared and cast liners subjected to firing tests 
in caliber .50 aircraft machine gun barrels. The re¬ 
sults are described in Section 19.8.2. The experi¬ 
mental program of Division 1 was terminated before 
the full potentialities of these materials could be 
evaluated. The following observations seem war¬ 
ranted by the data available on thermal and mechan¬ 
ical properties of these alloys at both room tempera¬ 
ture and elevated temperatures combined with the 
limited firing results: 

1. Some of these alloys besides the stellites show 
promise for use as liners in machine gun barrels. Fur¬ 
ther test and development (particularly the effect of 
heat-treatment on the hardness and properties) 
should be given to Refractaloy No. 70 92 and to other 
Refractaloy-type alloys and firing tests should be 
made on liners of N-155 alloy with about 0.25 or 
0.35% carbon. 

2. Chromium and cobalt, particularly a combina¬ 
tion of both in these alloys, enhances their resistance 
to chemical attack by powder gases. 

3. None of these alloys has a sufficiently high melt¬ 
ing point to be used as a bore-surface material in a 
gun under hypervelocity conditions where melting 
plays an important part in erosion. Most of them 
have fusion ranges between 1200 and 1300 C. It 
might be possible to utilize the excellent hot-hardness 
and ductility of some of these materials as liners, 
either under hypervelocity conditions or in machine 
guns under severe firing conditions at an increased 
velocity obtained by use of increased powder charges 
of single-base powder or by the use of double-base 
powder, if their rifled bores could be protected from 
melting by the use of a high-melting erosion resistant 
plate (such as chromium plate) provided that such 
plate can be made adherent. This possibility should 
be explored. 

4. Trials are warranted of liners of some of these 
materials in medium caliber guns at conventional 


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352 


SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 


velocities under more severe firing schedules or at 
regular schedules to obtain longer life. Since some of 
the alloys are amenable to fabrication as wrought 
ingots or tubes, they might be more easily applied to 
such guns than cast alloys. 

5. Because of their excellent strength at high tem¬ 
peratures as compared to gun steels, some of the more 
ductile of these alloys might be considered as gun 
barrel materials in order to decrease the weight of the 
gun tube and also to enhance the performance of an 
erosion resistant liner of a high-melting alloy (for 
example, a hardened molybdenum or a chromium- 
base alloy liner) or a high-melting erosion resistant 
plate (for example chromium electroplate). 

16,411 Steels and Other Iron Alloys 

Summary 

No steels or special high-iron alloys tested have 
shown any outstanding promise as bore-surface ma¬ 
terials. The results of laboratory and firing tests de¬ 
scribed in Section 16.3 and the results of studies of 



Figure 9. Cyclops KL liner fired 645 rounds with 
standard hall bullets M2 in caliber .50 heavy machine 
gun barrel. Cross section in. beyond the origin of 
rifling. Severe gas erosion obliterated the rifling. The 
light-gray layer at the surface is austenite containing 
undissolved carbides; below it is a tempered steel layer 
containing a large number of precipitated carbides. 
Etched with nital; 100X. (This figure has been taken 
from a progress report on Contract OEMsr-537.) 


the nature and mechanism of the erosion process and 
action of propellants on steel gun bores described in 
Parts III and IV, show that these alloys in general 
lack thermochemical resistance to powder gas ero¬ 
sion. Laboratory tests and firing tests of specific high- 
iron alloys as gun liners in machine gun barrels 
showed chemical attack by hot powder gases. Even 
when the hot-hardness was sufficient to prevent de¬ 
formation of the rifling by swaging impact of the 
bullet and the ductility was adequate to prevent brit¬ 
tle failure, chemical attack, and thermal transfor¬ 
mations at the bore surface (usually both) caused 
ultimate and usually early failure. As an example, a 
cross section of a Cyclops KL steel liner is shown in 
Figure 9. 

Two aircraft engine valve steels (Silchrome XCR 
and XB), which show somewhat better erosion resist¬ 
ance than ordinary gun steel, might be used as a 
barrel material for caliber .30 machine gun barrels 
under mild firing conditions or as liners or linings in 
large guns at moderate rates of fire where swaging of 
the rifling may be a minor factor in performance. 

Nickel and Cobalt Steels 

A liner of unhardened stainless steel (25 Ni, 12 Cr) 
in a caliber .30 machine gun barrel showed perform¬ 
ance 78 inferior to gun steel and evidence of thermo¬ 
chemical attack. A firing test 80 of a liner of a hard¬ 
ened special stainless steel (TEW alloy with 30 Ni, 
20 Cr, 4 Mo, 4 W, 2 Ta, 0.1 C, balance iron) in a 
caliber .50 heavy machine gun barrel showed that 
this alloy was subject to thermochemical attack by 
hot powder gases. On the other hand, a similar alloy 
(Refractaloy No. 70) 92 with 30% cobalt present and 
lower nickel and iron contents showed promise in this 
weapon. (See Section 16.4.10.) 

An iron-cobalt alloy (50 Fe, 48 Co, 2 V) in the 
erosion vent-plug test with %-in. vent showed 27 ero¬ 
sion resistance comparable to gun steels for the first 
few rounds and then weight losses increased more 
rapidly than with gun steels. It has good hot-hard¬ 
ness 59 up to 600 C. No firing tests in guns were made 
on this alloy. 

Tungsten Steels 

Liners of 18-4-1 high-speed steel (18 W, 4 Cr, 1 V, 
0.4 Mo, 0.75 C), Voland No. 2 hot-die steel (9 W, 
2.5 Cr, 2.25 Co, 0.34 V, 0.39 C) and Cyclops KL 
hot-die steel (7 W, 7 Cr, 0.34 C) were subjected to 


CONFIDENTIAL 




EVALUATION OF LINER AND COATING MATERIALS 


353 


firing tests 80 in caliber .50 heavy machine gun barrels 
and in addition to brittle failure all showed chemical 
interaction with powder gases and formation of an 
altered surface layer and hence have no promise as 
bore-surface materials. 

Hot-hard steels such as the hot-working die steels 
do not retain their hardness at elevated temperatures 
much better than regular WD 4150 gun steel, if they 
are heat-treated to a hardness of 30 to 35 Rockwell C. 
At this hardness their machineability is fairly good. 
When heat-treated to 50 Rockwell C, some of these 
alloys retain their hardness remarkably well up to 
about 700 C, but then they are not readily machine- 
able and also ductility and impact resistance are poor, 
with the result that machine gun liners of these alloys 
showed brittle failure as well as evidence of chemical 
attack on firing. 

Iron-tungsten electroplates had good hot-hardness 
but were not resistant to powder gas erosion. 

Molybdenum Steels 

Monomorphous iron-base alloys (iron with molyb¬ 
denum up to 20% and very low in carbon), which can 
be readily hardened by precipitation or aging, when 
tested as vent plugs showed large weight losses as a 
result of chemical attack. 

A 3% molybdenum steel (0.13-0.18 C) was found 
to have better erosion resistance than WD 4150 gun 
steel with single-base powder but, with double-base 
powder, it resisted erosion no better than the gun 
steel. 

Silicon Steels 

As already mentioned in Section 16.3.2, powdered 
ferrosilicon was resistant to chemical attack by single¬ 
base powder and retarded oxidation of gun steel dur¬ 
ing firings with double-base powder. A liner of a 
silicon steel (4.7%) tested 76 in the caliber .50 erosion¬ 
testing gun disintegrated by brittle failure after a few 
rounds. Unfortunately all iron-silicon alloys are too 
brittle to be satisfactory as liners or coatings in gun 
tubes and in addition their melting points are too low 
for severe service. 

When tested as a liner 81 in the caliber .50 heavy 
machine gun barrel, an aircraft engine valve steel 
Silchrome XCR (24 Cr, 5 Ni, 3 Mo, 0.45 C) 
showed much better resistance to thermochemical 
attack by powder gases than regular WD 4150 gun 
steel. In the unhardened condition this steel had ex¬ 


cellent tensile properties, hot-hardness, and ductility, 
but in the hardened condition, elongation dropped 
from about 22 per cent to around 3 per cent and liners 
failed by cracking. Silchrome XCR is difficult to 
prepare clean and free from oxide or carbide segrega¬ 
tions which make it difficult to machine. Another 
similar valve steel Silchrome XB (19.5 Cr, 1.35 Ni, 
2.29 Si, 0.76 C) was also being machined into barrels 
for the caliber .30 machine gun. This steel was be¬ 
lieved to have good erosion resistance on the basis of 
vent-plug data. Termination of the experimental pro¬ 
gram of Division 1 prevented completion of caliber 
.30 machine gun barrels of these steels and their 
testing, which should be completed. 

Nitrided Steel 

Laboratory tests and firing tests showed that WD 
4150 gun steel and Nitralloy when hardened by nitrid¬ 
ing or induction hardening of the bore showed no 
better resistance to powder gas erosion than in the 
unhardened condition. The use of hardened steel 
bores protected by an erosion-resistant chromium 
plating to improve the performance of machine gun 
barrels is described later in this report (Chapter 23) 
and the use of special steels with better high tem¬ 
perature properties than gun steels to enhance the 
performance of erosion resistant liners or plates or 
combinations thereof is described later in this report. 
(Chapters 23 and 24.) 

16 412 Copper and Its Alloys 

Pure copper and high-copper alloys show no prom¬ 
ise as bore-surface materials in spite of the fact that 
these materials show excellent resistance to chemical 
attack by hot powder gases. Even though the melting 
point of copper is relatively low, because of its high 
specific heat and excellent thermal conductivity, the 
melting point of a copper bore surface may not be 
attained during firing except under very severe con¬ 
ditions. Copper and all hardened copper-base alloys 
(for example, copper-beryllium alloys, which have 
excellent room temperature properties) lack, by a 
large margin, the necessary strength and hardness to 
resist severe deformation by powder pressures and by 
the swaging impact of the projectile at the working 
temperature in guns. The unsuccessful attempts to 
utilize copper in gun liners as one component of pow¬ 
der metallurgy compacts of chromium or molyb¬ 
denum or tungsten powders with copper was previ- 


CONFIDENTIAL 




354 


SELECTION OF EROSION RESISTANT MATERIALS FOR GUN BORES 


ously described in this report in Sections 16.4.3, 
16.4.4, and 16.4.5, respectively. Copper failed in firing 
tests 76 as an electroplate or as an undercoat under 
chromium electroplates (see Sections 16.4.2 and 
20.2.4). The severe plastic deformation or even melt¬ 
ing of bullet jackets of gilding metal (a high-copper 
alloy) may be a limiting factor in the performance of 
improved machine gun barrels under very severe 
firing conditions as mentioned again later in this 
report (Chapter 23). A rifled liner of copper hardened 
by alloying with a small amount of chromium when 
tested 78 by firing in a caliber .30 machine gun showed 
no chemical attack, but severe deformation and wear 
obliterating the rifling. 

16 413 Surface Coatings and Bore Lubricants 

Attempts have been made to find a nonmetallic 
material that can be applied to a gun steel bore sur¬ 
face to protect it from erosion by the powder gases, 
or which at least will reduce the frictional component 
of erosion. The latter purpose has also been extended 
to include the case of chromium-plated bores. None 
of the materials tried has been successful. 

A liner of WD 4150 gun steel whose rifled bore sur¬ 
face had been specially treated in steam at 1000 F to 
yield an oxidized bore surface of magnetic oxide of 
iron showed 76 no improvement in performance over 
untreated gun steel in the caliber .50 erosion-testing 
gun. 

One of the means suggested for protecting the bore 
surface was by a renewable film applied to the surface 
as each round is fired. The material selected for trial 
was a fluorocarbon, which is a chemically inert com¬ 
pound of high boiling point developed by Division 9, 
NDRC. 

In order to determine the effect of a film of fluoro¬ 
carbon on the erosion of gun steel, a special type of 
bullet with a grooved base-cup sealing ring was used. 
This groove, together with a deep groove in front of 
the base-cup, was filled with the fluorocarbon in the 
form of a grease before seating the bullet in the cal¬ 
iber .50 erosion-testing gun. It turned out that the 
film of fluorocarbon helped seal erosion cracks and 
thus reduced the customary drop in muzzle velocity, 111 
but the degree of erosion was just as severe as that of 
a control barrel with no protective film. 76 

The bore of a WD 4150 gun steel barrel for the 


m This result was later applied to the obturation of pre¬ 
engraved projectiles, as described in Section 31.4.2. 


caliber .50 erosion-testing gun was coated with a 
Parco-Lubrite coating approximately thick. 

This coating consists chiefly of a mixture of iron and 
manganese phosphates. It was believed that the re¬ 
duced friction between the bore and the projectile 
would result in an increased velocity life. A firing test 
showed no improvement in performance as compared 
with an untreated steel barrel. 76 It should be noted 
from Section 31.4.5, however, that a decided im¬ 
provement in performance is obtained when the pre¬ 
engraved projectiles are Parco-Lubrized. 

In firing tests in the caliber .50 erosion-testing gun 
on barrels or liners with chromium-plated bores, the 
life of the chromium plate was prolonged and per¬ 
formance improved by the use of cadmium-plated 
bullets, as mentioned in Section 31.4.5. On the other 
hand, trials 81 with cadmium-plated AP bullets M2 in 
nitrided and chromium-plated caliber .50 aircraft 
machine gun barrels showed no improvement com¬ 
pared with similar barrels fired with the same bullets 
not plated. 

16-4,14 Other Metals and Alloys 

In erosion vent-plug tests (see Section 16.3.1) zinc, 
beryllium, aluminum, and titanium showed 27 such 
severe thermochemical attack by hot powder gases 
that no quantitative data could be obtained. Espe¬ 
cially in the test of titanium the evolution of much 
smoke indicated that the action might be largely 
chemical. 

Silver showed poor performance in these same tests. 
It is not known for sure whether the high weight 
losses in this case were a result of melting 48 or a com¬ 
bination of melting and chemical attack. 

Pure columbium showed thermochemical resistance 
comparable to gun steels in erosion vent-plug tests 27 
but when filings of this metal were mixed with the 
propellant and fired into an evacuated tube (see 
Section 16.3.2) interaction with powder gases was 
observed. The lack of agreement between these two 
results has not been explained. 

Erosion vent-plug tests 75 of an alloy of gold (80%) 
and palladium (20%) were undertaken as part of the ex¬ 
perimental program to evaluate the relative impor¬ 
tance of melting and chemical attack in erosion. (See 
also Section 15.4.1.) This alloy has thermal properties 
similar to those of gun steel but a slightly lower melt¬ 
ing point. When tested under conditions such that the 
melting points of both the gold-palladium and the 
steel were exceeded, the former eroded somewhat 


CONFIDENTIAL 






EVALUATION OF LINER AND COATING MATERIALS 


355 


more than the latter. Under milder conditions, how¬ 
ever, the order was reversed and the relative amounts 
of erosion indicated a much higher resistance of the 
gold-palladium alloy to chemical attack than of gun 
steel. 

Indirect evidence 49 of the resistance of the pure 
metals rhodium and iridium to chemical attack by 


powder gases at very high temperatures was obtained 
when filings of these metals were used in experiments 
to determine the adiabatic flame temperatures of 
propellants. When the melting point of the metal was 
above the flame temperature, the particles remained 
unaltered. If chemical attack had occurred the ob¬ 
served behavior would not have been realized. 




CONFIDENTIAL 



Chapter 17 


CHROMIUM AND CHROMIUM-BASE ALLOYS 

By Helen M. W atson a 


INTRODUCTION 

17,1,1 Scope of Investigation 

C hromium was one of the first metals considered 
by Division 1 in the search for a bore-surface 
material that would solve the erosion problem in 
guns. It had been found that it possessed resistance 
to powder-gas erosion and some very desirable phys¬ 
ical properties; but early experiments by Division 1 
indicated that in order to be useful as a gun-liner ma¬ 
terial it would have to be prepared in a form stronger 
and less brittle than then known. 

An extensive investigation on the preparation of 
pure chromium was undertaken for Division 1 at the 
Westinghouse Research Laboratories in cooperation 
with the Geophysical Laboratory, C. I. W., and the 
National Bureau of Standards. No noticeable im¬ 
provement in ductility could be obtained by control 
of purity, by addition of small amounts of alloying 
elements, or by hot-working. Hence it was concluded 
that it would be unprofitable to continue the effort to 
develop pure chromium (or a high-chromium alloy) 
for use as a gun liner. 

Later, some of the hot-hard chromium-base alloys 
that had been developed by the Climax Molybdenum 
Company for Division 18, NDRC, appeared to offer 
possibilities as liner materials. The Climax Company, 
through a contract supervised by Division 1, invest¬ 
igated such applications of the alloys. Two of the 
alloys (those with the following percentage composi¬ 
tions: chromium 60, iron 25, molybdenum 15; and 
chromium 60, iron 30, molybdenum 10) were found 
to be very promising for this purpose. The low duc¬ 
tility of these alloys requires that they be properly 
supported in the gun barrel, as brought out in Section 
26.5.1. In every other respect, particularly by reason 
of inertness to the erosive action of double-base pow¬ 
der, caliber .50 liners of chromium-base alloys have 
given superior performance in firing tests. Another 
desirable feature is the availability of chromium, 
compared with cobalt and molybdenum, for example. 
Furthermore, its cost is not prohibitive. 


a Technical Aide, Division 1, NDRC. (Present address: 
Department of Physics, The Catholic University of America.) 


Thus these chromium-base alloys represent one of 
the three main solutions obtained by Division 1 to 
the problem of an erosion-resistant bore-surface ma¬ 
terial. Their potentialities as liners appear to be two¬ 
fold : for machine guns at high rates of fire, almost with¬ 
out regard to velocity, and for hypervelocity medium- 
caliber guns. Their further development, especially 
with respect to increase of ductility, was transferred 
by Division 1 to the Union Carbide and Carbon Re¬ 
search Laboratories, and then was continued there by 
the Army Ordnance Department after Division l’s 
contracts had been terminated. 

17,1,2 Preliminary Survey 

Examination of Chromium-Plated Naval Guns 

It has been the practice for nearly 20 years to apply 
a plating of chromium to the bore surfaces of Naval 
guns as a protection against corrosion during periods 
of inactivity. Microscopic examination of an eroded 
5 in./38-cal. chromium-plated gun revealed vestigial 
areas of the plate that were not perceptibly altered, 
even though the surrounding regions were severely 
eroded, 49 as is shown in Figure 2 of Chapter 20. 

The plate had failed by cracking, which permitted 
the powder gases to attack the underlying steel. How¬ 
ever, it was evident that the plate itself had with¬ 
stood the erosive action of the gases, as well as the 
wear caused by the projectile. This indication of the 
erosion resistant properties of chromium led to at¬ 
tempts to prepare a liner of pure chromium (Section 
17.2.2). 

Erosion Vent-Plug Tests 

Another basis for the belief that chromium might 
be suitable for an erosion-resistant bore surface was 
afforded by erosion vent-plug tests (Section 11.2.3). 
Such tests indicated that chromium is one of the few 
pure metals that are resistant to powder-gas ero¬ 
sion. 27 When tested with double-base powder in the 
PiQ-in. diameter erosion vent-plug apparatus, samples 
of a 50-chromiun^ 50-iron alloy eroded more than did 
gun steel, and the erosion rate of a 70-chromium, 30- 


356 


CONFIDENTIAL 



INTRODUCTION 


357 



Figure 1. Induction-heated vacuum furnace for melting chromium and high-chromium alloys. (This figure has 
appeared as Figure 2 in NDRC Report A-411.) 


CONFIDENTIAL 






































































































358 


CHROMIUM AND CHROMIUM-BASE ALLOYS 


tungsten alloy was similar to that of gun steel. Later 
the vent-plug tests were decreased in severity and 
their results more nearly approached those obtaining 
in guns (Section 11.2.3). The latter tests indicated 
that high-chromium ternary alloys of the chromium- 
iron-molybdenum and chromium-iron-tungsten series 
displayed very good erosion resistance. 75 

During the development of these chromium-base 
alloys, b the portions of 13 binary and 9 ternary alloy 
systems that contained more than 50 per cent chro¬ 
mium were surveyed. 167 - 168 Two of these systems, the 
chromium-iron-molybdenum and the chromium-iron- 
tungsten, were studied in detail. Their selection was 
based on stress-rupture tests at 870 C (1600 F), in 
which alloys belonging to the two systems displayed 
the highest strengths with measurable ductility. The 
physical properties of such alloys suggested the de¬ 
sirability of applying them to the gun-liner problem. 

172 METHODS OF PREPARATION 
17,2,1 Introduction 

Similar methods were employed for the preparation 
of both the pure chromium and chromium-base al¬ 
loys, namely, melting and casting high-purity stock 
under controlled conditions. Since brittleness is char¬ 
acteristic of these materials, care had to be taken that 
during the melting and casting process they would be 
kept as free as possible of impurities that would de¬ 
crease their ductility. It was found that by melting 
and casting them in vacuum, the oxide-, nitride-, and 
carbon-contents, which tend to increase their brittle¬ 
ness, were kept at a minimum. 

17 2 2 Chromium 

Vacuum Melting 

A specially designed induction furnace 83 (Figure 1) 
was constructed after it was found that unsatisfac¬ 
tory melts were formed when the entire charge of 
chromium chips was placed in a crucible and heated 
inductively. Chromium metal was “shoveled” gradu¬ 
ally from a charging side-arm down the charging chute 


b These alloys had been developed by the Climax Molyb¬ 
denum Company under an OSRD contract which was super¬ 
vised by the War Metallurgy Committee of the National 
Academy of Sciences in a search for heat-resistant metals for 
gas turbine blades. See Chapter 5 of the Summary Technical 
Report of Division 18. 


into a crucible mounted in the furnace tube, at such a 
rate as to avoid the formation of a crust on the sur¬ 
face of the melt. When melting had been completed, 
the power was adjusted and, while still in the fur¬ 
nace, the melt was cooled progressively from the bot¬ 
tom upward. 

Melting was done in a nearly complete vacuum at 
first, but it was observed that chromium vapor tended 
to condense on the cooler surfaces of the melting 
chamber and block openings that had been provided 
for sighting and charging. These difficulties were 
minimized when purified hydrogen or argon, at pres¬ 
sures of a few centimeters of mercury, was introduced 
into the vacuum chamber during melting. It was im¬ 
portant that the melting atmosphere be free of nitro¬ 
gen, since this element contaminated the melt, as 
shown in Figure 2, which is to be compared with 
Figures 3 and 4. 

The selection of a nonreactive crucible material 
was made especially difficult by the relatively high 
melting point (between 1900 and 2000 C) of chro¬ 
mium. Of the various crucibles tested, those made 
from pure aluminum oxide, without the use of any 
siliceous bonding material, proved to be the most 
satisfactory. 

Microexamination of chromium ingots produced in 
the early experiments revealed the presence of numer¬ 
ous inclusions (Figure 3) which were thought to be 



Figure 2. The structure of chromium contaminated 
by nitrogen. The nitrogen-bearing constituent is in¬ 
dicated at A. 250X. (This figure has appeared as Figure 
3 in NDRC Report A-411.) 


CONFIDENTIAL 






METHODS OF PREPARATION 


359 


chromium oxide (Cr 2 0 3 ). These presumably resulted 
from either (1) the oxygen known to be present in the 
electrodeposited chromium used as the source ma¬ 
terial, or (2) the oxygen picked up from residual gases 
liberated from the refractories during melting, or both. 
Zirconium metal was a very effective deoxidizer when 
added to the melt, as shown by a comparison of the 
structure shown in Figure 4 with that in Figure 3. In¬ 
gots weighing about 300 g were produced. 


.‘V 

/ . m 

• 

. *•'■*>* 

: . 

G / 

, ■* ' . -1 

A* * 

£> ■ 4 

• . •• . * 

« k 

■ ; _ 0 9 

* At 

® " . * 


, * 

Hi 


Figure 3. A typical section of a pure chromium ingot 
that had been cooled in the furnace after having been 
melted in vacuum without a deoxidizer. Oxide inclu¬ 
sions appear along the grain boundaries. 250X. (This 
figure has appeared as Figure 4 in NDRC Report A-411.) 

The chromium used for melting stock was prepared 
by electrodeposition. It was very pure except for its 
oxygen content, which was equivalent to 0.1 to 1.5% 
of chromium oxide (Cr 2 03 ). In an attempt to remove 
the oxygen and thereby effect an increase in ductility, 
a series of “beneficiation” experiments was carried 
out. Samples of electrodeposited chromium were 
heated at temperatures up to their melting points in 
a current of hydrogen purified by diffusion through 
palladium. This treatment removed most of the oxy¬ 
gen originally present in the chromium, and improved 
its cold ductility to a slight extent. 

Similar heating in vacuum produced needle-like de¬ 
posits of condensed chromium, some of which could 
be bent repeatedly through large angles or, in a few 
cases, rolled to form thin, irregular plates that would 


stand some plastic deformation without fracture. The 
behavior of such crystals was observed microscopi¬ 
cally, and it was concluded that whether fracture 
would result upon the application of compressive 
forces to the crystal depended upon whether the 
forces were applied normal to an octahedral plane 
(no cracking) or to a cubic plane (cracking); and that 
therefore no significant degree of cold ductility could 
be expected in polycrystalline masses of pure chro- 


V *,v 



Figure 4. A typical section of a pure chromium ingot 
that had been cooled in the furnace after having been 
melted under a low pressure of argon and deoxidized 
with zirconium. 250X. (This figure has appeared as 
Figure 5 in NDRC Report A-411.) 

mium where the orientation of individual crystallites 
is random. 

Furnace-cooled ingots nearly free of oxide inclu¬ 
sions and otherwise of high purity were obtained 
starting with “beneficiated” chromium; but their 
ductility, hardness, and tensile properties were so poor 
that it was not considered practicable to attempt to 
make gun liners from such material. 

Vacuum Casting 

All the furnace-cooled ingots had very coarse grain 
sizes. It was considered that if chromium having a 
finer grain structure than is possible with furnace 
cooling could be prepared, it might possess sufficient 
ductility. The furnace was redesigned to include a 


CONFIDENTIAL 






360 


CHROMIUM AND CHROMIUM-BASE ALLOYS 


tilting mechanism which permitted the transfer of 
the molten metal from the crucible to a steel mold 
without a change in the atmosphere. The two castings 
of zirconium-deoxidized pure chromium made in this 
way were quite brittle, and not very strong. They 
were remarkable for their highly developed radial 
crystallization patterns (shown in Figure 5) and for 
the soundness of the metal, except for round shrink¬ 
age cavities which extended almost the entire length 
of the ingot. 



Figure 5. A vacuum-cast ingot of pure chromium 
deoxidized with zirconium. It had a well-developed 
radial crystallization pattern. (This figure has appeared 
as Figure 9(b) in NDRC Report A-411.) 

The chill-cast specimens showed no improvement 
in malleability at room temperatures compared with 
those cooled in the furnace. Thereupon the attempts 
to produce a ductile, pure chromium in appreciable 
amounts were discontinued. 

Thermal Decomposition of Iodide 

A supplementary investigation was undertaken to 
prepare crystalline metallic chromium of an extremely 
high order of purity by a method that would minimize 
the possibility of oxygen contamination. A small 
amount of crystalline chromium was prepared by 
thermal decomposition of chromous iodide (Cr 2 I 2 ) on 
an electrically heated tungsten filament. 8 Consider¬ 
able cold ductility was exhibited by these crystals, 
but the method was considered impracticable for 
large amounts of the metal. Microexamination re¬ 
vealed these crystals to be cubic, and to show octahe¬ 
dral faces. The same observations were made con¬ 
cerning their anisotropy with respect to malleability 


as were made for the needle-like crystals that resulted 
from beneficiation experiments described previously. 

Electrodeposition on Brass 0 

An attempt was made, early in the investigation, 
to prepare chromium liners by electrodepositing chro¬ 
mium on brass tubes to a thickness of }/$ inch and 
then dissolving out the brass. 49 A tube deposited at 
75 C with a current density of 100 amp/dm 2 was 
found to be free of initial cracks, but was exceedingly 
brittle. Another tube, which had been softened and 
made more ductile by annealing, was shrunk into a 
steel tube and rifled. When it was tested as a liner in 
the heavy barrel of a caliber .50 Browning machine 
gun (Section 11.2.2), many longitudinal cracks devel¬ 
oped after the firing of only a few rounds. No im¬ 
provement resulted from further experimental work, 
and it was concluded at that time that mechanical 
working would be necessary before electrodeposited 
chromium would have mechanical properties suitable 
for use as a liner. However, later attempts to improve 
the properties of chromium by hot-working led to the 
conclusion that no noticeable benefit could be effected 
(Section 17.3.2). 

17,2,3 Chromium-Base Alloys 

Vacuum Casting 

Apparatus used 87 for vacuum-casting chromium- 
base alloys is shown in Figure 6. As was the case 
in preparing chromium, particular care had to be 
taken in choosing the crucible used for melting, to in¬ 
sure that there would be no reaction with the molten 
material. Alundum crucibles were tried, but it was 
found that they melted at the high temperatures 
needed to cast the melt into a mold. Beryllia or zir- 
conia crucibles, wrapped with molybdenum sheet as a 
resistor, proved to be the most satisfactory. 

The alloys were melted by induction heating, and 
then were transferred through a hole in the bottom of 
the crucible to a copper casting mold. The metal mold 
chilled the melt rapidly, and thus produced fine pri¬ 
mary grains. However, it was found to be important 
that the alloy be chilled from the melt at the max¬ 
imum rate that did not cause cracking or induce un¬ 
favorable residual stresses. If chilling was too rapid, 


c See Chapter 20 for a discussion of electrodeposition on 
gun-bore surfaces. 


CONFIDENTIAL 









METHODS OF PREPARATION 


361 


OBSERVATION WINDOW 



Figure 6. Vacuum-casting apparatus for preparation of centrifugally cast caliber .50 gun liners made of chromium 
base alloys. (Figure 1 in NDRC Report A-415.) 


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362 


CHROMIUM AND CHROMIUM-BASE ALLOYS 



Figure 7. Induction-heated apparatus for centrifugal in-melting of chromium-base alloy tubes. (Figure 1 in NDRC 
Report A-410.) 

The Micarta furnace tube I contained the powdered charge in a thin-walled carrier tube surrounded by granular refractory in which longitudinal 
cooling tubes are imbedded. The tapered ends of the furnace tube fit against gasketed chucks J, carried by hollow shafts K, running in bear¬ 
ings in the head-stock L and tail-stock M of the lathe N. Cooling water is admitted through the inlet tube G. Coil springs O serve to push the 
movable bearing shell P longitudinally toward the head-stock and thus maintain a tight connection between I and J. The inductor coil Q is 
moved longitudinally along the ways of the lathe by the lead-screw R, which is driven by an adjustable-speed motor. 


large temperature gradients, which often caused 
cracks, resulted. 

The copper casting mold was rotated at speeds of 
from 4,500 to 5,500 rpm, so that tubular castings 
were produced with parabolic inside surfaces. The 
diameter of the bore varied from in. at the top to 
34 in- at the bottom. About 20 min after they had 
been poured, the castings were removed from the 
mold and were placed in an air furnace at 870 C 
(1600 F), where they were held for a period between 
30 and 120 min for relief of casting stresses. The cast¬ 
ings were then ready for machining. 

This apparatus was capable of melting charges up 
to 434 lb. An improved apparatus, designed to melt a 
50-lb charge, was constructed but had not been as¬ 
sembled when the investigation under NDRC auspices 
was discontinued. d In addition to the larger charge, 
the improved design provided for a continuous opera¬ 
tion of the apparatus, brought about by the addition 


d The improved apparatus was constructed for Climax 
Molybdenum Company, and at the termination of its contract 
the equipment was transferred to the Union Carbide and Car¬ 
bon Research Laboratories where experiments were continued. 


of an air lock between the melting and the casting 
chambers. A tilting mechanism for the crucible, which 
enabled the melt to be poured over its lip into the 
casting mold, eliminated bottom pouring. 

Centrifugal In-Melting 

A new method of preparing alloy tubes, termed 
progressive centrifugal in-melting, 82 was tried at West- 
inghouse Research Laboratories with an alloy of the 
approximate composition 60-chromium, 35-iron, 5- 
molybdenum. The melting took place inside a Mi¬ 
carta tube (about 4 in. in diameter and 36 in. long) 
lined with granular magnesium oxide or aluminum 
oxide, in the center of which the charge was intro¬ 
duced in the form of metal powder. The tube was 
evacuated and then rotated at about 1,200 to 1,400 
rpm while an inductor coil was moved slowly from 
one end to the other. In this way successive small por¬ 
tions of the charge were melted and then subsequently 
frozen while under the influence of centrifugal force. 
The Micarta tube was protected from the heat of the 
core by a “squirrel cage,” consisting of a system of 
longitudinal water-cooled copper tubes. The Micarta 


CONFIDENTIAL 



















PHYSICAL PROPERTIES 


363 


Table 1 . Physical properties of chromium and chromium-base alloys compared with those of gun steel. 


Property 

Chromium i 

60 Cr—25 Fe—15 Mo 

60 Cr—30 Fe- 

10 Mo Gun steei* 

Density (g/cc) 

7.14 

7.6 


7.83 

Melting point (C) 

1950 ± 50 

ca. 1700 

ca. 1650 

1450 

Specific heat (cal/g/°C) 

0.12 



0.11 

Latent heat of fusion (g-cal/°C @ 15 C) 

31.75 



Thermal conductivity (cal^/cm/sec/°C) 

0.165 (15 C) 

0.0305 (30 C) 


0.102 (50 C) 

Coefficient of thermal expansion ( /°C) 

8.1 X 10 ® 

7.7 X 10~ 6 



(© 20-300 C) 

(© 25-315 C) 





8.8 X lO" 6 

10.8 X 10~ 6 

13.6 X 10" 6 



(@ 25-650 C) 

(@ 25-595 C) 

(@ 20-500 C) 

Hardness @ 20 C (68 F) 

193 

480 ± 10 

420 + 15 

290 1 

(VPN) @ 800 C 




36 f 

@ 870 C (1600 F) 

142 

198 

149 

Modulus of elasticity (psi) 


33 X ion 

33 x ion 

29-30 X 10 6 

Tensile strength (psi)§ 


ca. 100,000 

135,000 (70 F) 



(@ 1350 F) 


45,000 (1200 F) 

Stress-rupture strength @ 870 C; (1600 F) 




Stress (psi) 

20,000 

20,000 

20,000 


Time for rupture (min) 

1 

3,000 

240 


Elongation (%) 

3.5 

13.0 

26.0 


Reduction of area (%) 

3.7 

12.2 

24.6 



* Oil-quenched and tempered; approximating the composition of SAE 4150 steel, 
t From Table 2 of Chapter 19. 

J Average value obtained from preliminary tests. 

§ For other tensile properties of steel see Table 4 of Chapter 19. 


furnace tube I and the inductor coil Q are shown in 
Figure 7, together with the means employed for oper¬ 
ating the furnace. 

The experiments had not been carried beyond the 
exploratory stage by the time the project was ter¬ 
minated. Several sound tubular castings, free from po¬ 
rosity, were obtained; but no metallographic exami¬ 
nation was made to find out whether there was much 
grain-boundary contamination by oxides, carbides, 
and nitrides. Inasmuch as the method has inherent 
advantages over casting in the case of large-diameter 
tubes, it would appear to be worth further investiga¬ 
tion as a means of making gun liners of chromium- 
base alloys. 

17,2-4 Chromium Impregnations 

The brilliant cleanliness of beneficiated chromium 
(Section 17.2.2) suggested an investigation of metal- 
bonded chromium compacts. Several compacts were 
prepared with copper, nickel, or palladium as impreg- 
nants. It was hoped that the resulting composite 
materials would have the basic properties of chromi¬ 
um, together with the useful mechanical properties of 
the impregnant. 

Tests of copper-bonded chromium showed a ten¬ 
sile strength of 100,000 psi (that of annealed copper is 
36,000 psi), but with very poor ductility. When a 


sample of the compact was tested as an erosion vent- 
plug 49 it stood up better than gun steel during the 
first three rounds, but cracked badly on the fourth. 

173 PHYSICAL PROPERTIES 6 

17,3,1 General Resume 

Chromium 

Chromium is a very hard gray metal, resembling 
iron. Its principal ore is chromite, a complex oxide of 
iron and chromium containing one-third to one-half 
chromium oxide (Cr 2 03 ). The metal is prepared by re¬ 
duction of the oxide by aluminum. For some purposes 
the metal is then purified by electrolysis. Some of the 
properties of chromium are listed in Table 1. 

Among the factors that prompted the investigation 
of chromium as a material for gun liners was its high 
melting point. Only nine other metals have higher 
melting points: rhodium, masurium, iridium, rutheni¬ 
um, molybdenum, osmium, tantalum, rhenium, and 
tungsten. The cost of chromium in the pure state is 
relatively high, not because of any scarcity of ore, but 
because of the expense in reducing the ore. 


e The data given in this section are taken from several 

sources 83,87,167,168,504,524 


CONFIDENTIAL 































364 


CHROMIUM AND CHROMIUM-BASE ALLOYS 


Chromium-Base Alloys 

The chromium-base alloys that were considered 
worth testing as gun liners were those containing from 
50 to 60% chromium, from 25 to 45% iron, and from 
5 to 15% molybdenum or tungsten. Since alloys 
with molybdenum displayed greater ductility than 
those containing tungsten, the latter were soon ruled 
out, and attention was centered on the chromium- 
iron-molybdenum system. The two most promising of 
this ternary alloy system proved to be 60-chromium, 
25-iron, 15-molybdenum and 60-chromium, 30-iron, 
10-molybdenum. 

The chromium’s principal contribution was chem¬ 
ical inertness and high melting point; the iron’s, duc¬ 
tility ; and the molybdenum’s, erosion resistance and 
strength. Since iron is similar to chromium, it does 
not greatly reduce the high melting point of the latter. 
Moreover, its presence permits the use of ferrochrom- 
ium as source material of the chromium. 



Figure 8. Typical structure of a chromium-base alloy 
gun liner before firing. (This figure accompanied the 
manuscript of NDRC Report A-415, but it was not 
reproduced in that report.) 

Centrifugal cast alloy stress-relieved 2 hr at 870 C, for liner 
L-123, assembly CX-30. C composition: 60Cr-30Fe-10Mo contain¬ 
ing 0.03% C, hardness 407 VPN. Estimated percentage: of carbide, 
0.10; of oxide, 0.05. ASTM grain size: 1-3. Electrolytic etch in 10% 
oxalic acid. 100X. 


Since chromium, iron, and molybdenum all crystal¬ 
lize in the body-centered-cubic arrangement, it was 
expected that their alloys in the composition range 
indicated above would possess similar crystalline 
structures. Nothing that would indicate otherwise 
has yet been observed. Evidence has been found of an 
intermetallic compound that i^ precipitated from the 
chromium-rich solid solution. It is believed that the 
compound may be FeCr. When such precipitation oc¬ 
curs, the hardness of the alloy is increased. It is not orig¬ 
inally present in chill-cast alloys, but it precipitates on 
holding the alloy for several hours at 870 C (1600 F). 

The melting points of the alloys in the chromium- 
iron-molybdenum system are not nearly so low as 
that of steel provided that the percentages of the im¬ 
purities (carbon, oxygen, nitrogen, and silicon) are 
low. 

Included in Table 1 are some of the physical prop¬ 
erties of the two most promising (from the stand¬ 
point of hypervelocity-gun liner materials) of the 
chromium-base alloys investigated: 60-chromium, 
25-iron, 15-molybdenum and 60-chromium, 30-iron, 
10-molybdenum. Data were not available to complete 
the table, since only a beginning has been made in the 
investigation of the properties of chromium-base 
alloys. In some cases several values were available, 
but the one given in the table may be considered rep¬ 
resentative. Figure 8 shows the typical structure of 
a chromium-base alloy gun liner before firing. 

17 3 2 Efforts to Improve Ductility 
by Hot-Working 

Chromium 

The principal drawbacks to the use of chromium 
liners for gun barrels are its lack of ductility and its 
low strength. Since the former failing is the more 
serious, an investigation 83 was carried out to deter¬ 
mine whether plastic working of chromium at an 
elevated temperature would so alter the structure of 
the metal as to bring about some degree of plasticity 
at ordinary temperatures. 

The test assembly consisted of a tube of annealed, 
electrodeposited chromium fitted into a hollow bar of 
SAE 4140 steel. Connected to the assembly was a 
copper tube leading to a hydrogen supply, so that a 
flow of hydrogen was maintained through the small 
clearances between the inner and the outer parts of 
the test piece. The outer part of the assembly was 
heated in a hydrogen-atmosphere furnace and was 


CONFIDENTIAL 




FIRING TESTS OF LINERS 


365 


swaged, while hot. This action elongated both the 
steel and the chromium. 

Tests were carried out at several temperatures, 
ranging from 750 C to 1400 C. With respect to plastic 
deformation of the chromium, 1200 C was the most 
favorable temperature. It was observed that at 750 C 
the chromium was shattered. Only at 1400 C was there 
any bond formed between the steel and the chromium. 
When the latter was freed it was found to be still 
quite brittle, and penetrated by longitudinal or helical 
radially-disposed cracks. 

There was no noticeable improvement in cold duc¬ 
tility of the chromium from the hot-working. These 
experiments demonstrated quite forcibly the relatively 
great hot-hardness of chromium compared with that 
of steel. The chromium had behaved as a somewhat 
deformable body embedded in another softer and 
quite plastic one, the steel having flowed past the 
chromium and dragged it along. 

Chromium-Base Alloys 

Hot-working experiments 87 were also performed 
with some chromium-base alloys. The grain size of 
the alloy is a significant factor in its performance as a 
gun liner, and forging increases the alloy’s ductility 
by refining the coarse, as-cast grains. Some alloy in¬ 
gots, from 8 to 9 in. in length, having a 1%-in. diam¬ 
eter at the bottom and l^g-in. at the top, were in¬ 
serted in cylindrical hollows in swages made from 
SAE 4140 steel and from 3%-molybdenum iron. An 
inductively-heated vacuum furnace was used. 

The ease with which the alloy could be swaged de¬ 
creased as the iron-molybdenum ratio decreased. In¬ 
gots of the composition 50Cr-45Fe-5Mo were readily 
hot-worked at temperatures of from 1260 to 1340 C. 
A gun liner was forged from an ingot of this composi¬ 
tion. At a slightly higher temperature (1370 to 1435 C), 
ingots of the 60Cr-35Fe-5Mo composition were 
easily forged, with very few surface cracks. Consider¬ 
ably more difficulty was encountered with 60Cr- 
30Fe-10Mo ingots for which a forging temperature of 
from 1435 to 1480 C was required. Circumferential 
cracks, apparently caused by longitudinal tensile 
stresses set up at the corners of the swage, were ob¬ 
served. 

17 3 3 Fabrication of Chromium-Base Alloys 

Because of their higher hardness and greater tend¬ 
ency to tear out, chromium-base alloys whose chro¬ 


mium content is about 60% are, on the whole, less 
machineable than steel and Stellite No. 21. The ma- 
chineability was found to be a function of the iron- 
molybdenum ratio. Alloys of the composition 60Cr- 
25Fe-15Mo and 60Cr-30Fe-10Mo, in which the ratio 
is at least 5 to 3, can be machined with high-speed 
steel tools. Alloys having the ratio less than 5 to 3 but 
greater than 3 to 5 require cemented-carbide cutting 
tools, or grinding. Alloys with the ratio less than 3 to 
5 can be fabricated only by grinding. 524 

Because of limitations of time and personnel, an ex¬ 
haustive study of the best method of machining and 
casting gun liners of chromium-base alloys has not 
yet been undertaken. The method emp^ed during 
Division l’s investigation was based on fundamental 
principles, augmented by techniques that developed 
as the work progressed. The method 87 that evolved 
was practicable; although more desirable techniques 
no doubt would result from a search for the optimum 
tools and operations. 

The danger of nitrogen and oxygen contamination, 
with resulting increase in brittleness, prevents the 
welding of chromium-base alloys by the techniques 
applied to cobalt-, iron-, and nickel-base alloys. Pre¬ 
liminary experiments indicate that an inert atmos¬ 
phere is necessary, and that the welding rod should be 
prepared from certain chromium-base alloys. 168 

A very decided advantage of chromium-base alloys 
over molybdenum as a gun-liner material is that they 
can be cast as seamless tubes, or perhaps be prepared 
by centrifugal in-melting (Section 17.2.3). 

17 4 FIRING TESTS OF LINERS 87 
17,4,1 Introduction 

Liners of nine different alloy compositions were 
fired in the three types of caliber .50 testing guns de¬ 
scribed in Sections 11.2.1 and 11.2.2: the erosion-test¬ 
ing gun, the heavy barrel machine gun, and the air¬ 
craft machine gun. A strict comparison of the results 
of the 33 firing tests is not possible, since the attempts 
to improve the properties of the alloys paralleled ex¬ 
periments on methods of supporting the liner in the 
barrel. Therefore the later tests represented improve¬ 
ments in all these respects. It was determined that 
10% molybdenum was the minimum percentage for 
adequate erosion resistance and strength, and that 
.when properly supported in the gun barrel, liners of 
a range of compositions have sufficient ductility to 
withstand the powder pressure without cracking. 


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366 


CHROMIUM AND CHROMIUM-BASE ALLOYS 


17 4 2 Variation of Composition 

The first firing tests were of liners (in heavy bar¬ 
rels) made from alloys of the three compositions that 
had performed so well in erosion vent-plug tests (Sec¬ 
tion 17.1.2): 60Cr-25Fe-15 W, 55Cr-35Fe-10Mo, and 
60Cr-25Fe-15Mo. The results indicated that the al¬ 
loys possessed good erosion resistance and had suf¬ 
ficiently high hot-hardness and melting points. Crack¬ 
ing of the liners and constriction of the bores were 
the principal causes of failure. The 60Cr-25Fe-15Mo 
composition was the best of these three tested in 
the heavy barrel. This same conclusion was reached 
in tests in the erosion-testing gun of liners of the 
compositions 60Cr-25Fe-15Mo, 60Cr-25Fe-15W, and 
50Cr-45Fe-5Mo. 

Because of their greater ductility, alloys containing 
molybdenum were selected in preference to the tung¬ 
sten-bearing ones, and a series of tests to determine 
the most favorable composition was undertaken. This 
was determined to lie within the range represented by 
the compositions 60Cr-30Fe-10Mo and 60Cr-25Fe- 
15Mo on the basis of the following results of tests: 

1. The compositions 60Cr-35Fe-5Mo and 55Cr- 
43Fe-2Mo (the most ductile of the chromium-base 
alloys tested) had insufficient erosion resistance and 
strength. 

2. The compositions 55Cr-35Fe-10Mo and 55Cr- 
30Fe-15Mo had insufficient erosion resistance, al¬ 
though the hardness and strength were nearly satis¬ 
factory. The latter alloy had been heat-treated to a 
hardness of 613 VPN, as contrasted with 454 for the 
former. 

3. The composition 55Cr-25Fe-20Mo had about the 
same erosion resistance and slightly greater strength 
than 60Cr-25Fe-15Mo but its machineability was 
much less than that of the latter. 

17 4 3 Assembly Variations 

Since the two most promising alloy compositions 
for gun liners were found to be 60Cr-30Fe-10Mo and 
60Cr-25Fe-15Mo, liners prepared from the two alloys 
were given extensive firing tests. The former composi¬ 
tion possesses the greater ductility, but it represents 
the minimum percentage of molybdenum to impart 
sufficient erosion resistance and strength. 

The principal cause of failure of the liners during 
the firing tests was longitudinal cracking. A special 
study 87 was made of the stresses imposed on the liner. 
Formulas such as those given in Section 26.4.5 were 


used in computing the stresses produced by shrink¬ 
fitting liners in the various types of machine-gun 
assemblies. Cracking was prevented by imposing an 
initial hoop stress of from 90,000 to 100,000 psi on the 
liner. In order to achieve this degree of compression 
with the relatively thin aircraft barrel, a reinforcing 
sleeve had to be shrunk on the outside of the barrel. 
The sleeve was made of Timken No. 17-22A steel, 
which possesses greater high-temperature strength 
than ordinary gun steel. 

Because of the successful results obtained with 
two-stave molybdenum liners (Section 18.6), a two- 
stave assembly of the liner composition 60Cr-25Fe- 
15Mo was prepared and tested in the caliber .50 
erosion-testing gun with double-base powder. The 
results were disappointing in that longitudinal and 
circumferential cracks were observed in the fired 
liner. 

In a test to determine what benefits in firing per¬ 
formance would be effected by supporting the liner 
with a backing material possessing high elevated- 
temperature strength, an assembly was prepared in 
which a liner of the composition 60Cr-25Fe-15Mo 
was supported by a sleeve made from an alloy of the 
composition 60Cr-30Fe-10Mo, shrunk into the steel 
retainer. When tested in the caliber .50 erosion-test¬ 
ing gun with double-base powder, a total of 568 
rounds was fired. The test was discontinued because 
of excessive pressure rise caused by slight constriction 
and roughening of the surface. This represented the 
greatest number of rounds sustained by a chromium- 
base alloy assembly in the erosion-testing gun. 76 
Upon examination of the fired liner there was no 
evidence of attack by the powder gases, but cracking 
and pitting of the bore surface were observed. 77 Some 
of these effects can be seen in a cross section of the 
liner, 6 in. beyond the origin of rifling, shown in 
Figure 9. 

Although this test and the one using the stave-type 
of construction were not successful as far as method 
of insertion was concerned, they did prove the high 
order of resistance of these chromium-base alloys to 
powder gases, even those from double-base powder. 
The conditions of erosion were so severe in the caliber 
.50 erosion testing gun that a gun steel barrel was 
worn out after only 90 rounds. 

Preferred Method of Insertion 

The foregoing tests of different methods of inser¬ 
tion of chromium-base alloy liners in caliber .50 bar- 


CONFIDENTIAL 



CHROMIUM-BASE ALLOYS AS LINER MATERIALS 


367 



Figure 9. Cross section of 60Cr-25Fe-15Mo liner, 
6 in. beyond origin of rifling, after having been fired 
568 rounds with ball bullets and FNH-M2 powder. The 
surface of the liner had cracked and pitted along the 
grain boundaries and some pitting had developed. This 
liner had been supported in a caliber .50 erosion-testing 
gun tube by a sleeve of 60Cr-30Fe-10Mo. Etched with 
10% KOH—10% K 3 Fe(CN) 6 . 100X. (This figure has 
appeared as Figure 17 in NDRC Report A-405.) 


rels made clear the requirements for a successful 
assembly. These were that the liner be made cylin¬ 
drical, that it be restrained by a hoop stress of about 
100,000 psi, and that the steel of the barrel have 
sufficiently good high-temperature strength to main¬ 
tain this compression even after continued firing 
(Section 24.5). Furthermore it was desirable that the 
bore surface of the barrel ahead of the liner be pro¬ 
tected from the powder gases in order to utilize fully 
the great erosion resistance of the liner. f 

These requirements were met by inserting two 
liners each of alloys having the compositions that the 
previous tests had shown were most suitable (600- 
25Fe-15Mo and 60Cr-30Fe-10Mo) in 13-lb aircraft 
barrels made of Timken No. 17-22A steel (Section 
24.5.3), which had been chromium-plated ahead of 


f Many of the firing tests in the caliber .50 aircraft barrels 
had to be discontinued as a result of severe erosion of the steel 
barrel ahead of the liner, and others were continued long past 
the stage of “keyholing” caused by erosion there, just to test 
the endurance of the alloy. 


the liner recess with choked-muzzle chromium plate 
(Section 23.1.4). These barrel assemblies were tested 
in the caliber .50 aircraft machine gun by firing the 
CGL-350 schedule described in Section 23.1.3. After 
3,850 rounds the liners had not failed, but the test 
was discontinued because of termination of Division 
Fs contract with the Crane Company, 80 where the 
firing took place. Numerous surface checks were ob¬ 
served after the initial 350-round burst, and these 
irregularities in the surface had developed into large 
pits and cracks of indeterminate depth after the next 
group of 500 rounds. After that they showed little 
further change. 

The average “cold velocity” (Section 23.1.3) of the 
barrels containing chromium-base alloy liners in¬ 
creased gradually to a maximum after 3,350 rounds, 
which indicated that a slight progressive shrinkage of 
the bore near the origin of rifling took place. The “hot 
velocity” remained essentially unchanged during the 
test. The accuracy was poorer than that obtained in 
the firing of two barrel assemblies similarly prepared 
containing liners of No. 21 Stellite (Section 24.5.3); 
but no reason was discovered for the difference. The 
wear of the bore surface was less for all four chro¬ 
mium-base alloy liners than for the two stellite ones 
after the same number of rounds. This difference 
indicated that the chromium-base alloy liners might 
have lasted longer than the stellite ones, provided 
that they did not crack or shrink to such an extent 
that continued firing was impossible. 


ns POTENTIALITIES 

OF CHROMIUM-BASE ALLOYS 
AS LINER MATERIALS 

17,51 Extent of Suitability 

The properties of an ideal erosion resistant mate¬ 
rial are discussed in Section 16.2. They are reviewed 
here to see how well chromium-base alloys fulfill the 
requirements. 


Thermal Resistance 

The melting points of the chromium-base alloys 
have not been determined closely; but presumably 
they are not a great deal lower than that of pure 
chromium, since the effects of the iron and molyb¬ 
denum are in opposite directions. The specific heat 
is presumably about the same as that of steel whereas 
the thermal conductivity is only about one-third as 


CONFIDENTIAL 





368 


CHROMIUM AND CHROMIUM-BASE ALLOYS 


great. Hence the thermal factor pck in equation (13) 
of Chapter 5 is also about one-third as great as for 
gun steel, and consequently the increase in tempera¬ 
ture of a very thin layer at the surface after the firing 
of one round is considerably greater than for gun 
steel, perhaps as much as 50 per cent. 

The consequent thermal stress is also somewhat 
greater, although the effect of the greater tempera¬ 
ture increase is offset somewhat by the lower thermal 
expansion. This thermal stress applied to a material 
of low ductility does cause a slight checking of the 
surface. In this respect these alloys are inferior to 
stellite, when cool propellants are considered; but on 
the other hand, since the melting points are high 
enough to prevent the melting of a surface film when 
double-base powders are fired, as occurs with stellite, 
they may be said to be superior to stellite in overall 
thermal resistance, while inferior to molybdenum in 
this same respect. 

Chemical Resistance 

The alloys possess excellent resistance to the chem¬ 
ical action by the powder gases at elevated tempera¬ 
tures. 

Mechanical Properties 

It is in this category that the alloys display their 
serious weakness, poor ductility. In spite of this 
weakness, it has been possible to make successful 
caliber .50 liners, as described in Section 17.4.3. In 
other respects they meet the requirements of the 
ideal material: no large permanent change in dimen¬ 
sions (provided that casting strains are avoided), no 
permanent deformation or flowage, high strength at 
elevated temperature (especially in the 60Cr-25Fe- 
15Mo alloy), and thermal and elastic properties 
close to those of gun steel, except that the coefficient 
of thermal expansion is only about two-thirds as 
great. 

Other Properties 

The raw material for the alloys is available in 
sufficient quantity. The process used to prepare the 
experimental liners was satisfactory, and there is 
every indication that, with continued development, 
it ultimately can be adapted to commercial produc¬ 
tion. The alloys are readily machineable if proper 
tooling is used. 


17 5 2 Recommendations for 

Further Development 

% 

Firing Tests 

Since higher potential propellants offer a con¬ 
venient means of increasing the muzzle velocity of an 
existing gun, it is suggested that liners of the two 
chromium-base alloys that have proved to be most 
premising so far— 60Cr-25Fe-15Mo and 60Cr-30Fe- 
lOMo—be tried in the caliber .50 aircraft machine 
gun with the experimental double-base powder al¬ 
ready developed by the Army Ordnance Depart¬ 
ment. For such a test the liners might be inserted in 
a “finned” barrel, such as is described in Section 24.3.3, 
made of high-strength steel instead of a 13-lb barrel 
as was used in the tests described at the end of Sec¬ 
tion 17.4.3. Then the test could be made with the 
high-speed caliber .50 gun, M3, without sacrifice of 
cyclic rate. 

Improvement of Ductility 

Since brittleness is so serious a problem in the 
extensive application of these alloys, fundamental 
research should be continued to find a ductile com¬ 
position. Some experiments 87 at Union Carbide and 
Carbon Research Laboratories have indicated that 
the alloy 60Cr-25Fe-15Mo is very susceptible to work 
hardening. Since iron seems to improve the ductility 
more than any other alloying element, it may be 
possible to use a chromium-base alloy with higher 
iron content and obtain the necessary hardness by 
work-hardening. It is questionable, however, whether 
the thermochemical resistance of such an alloy would 
be sufficient even after hardening. 

Another possible way of obtaining a more ductile 
alloy is by the use of a fourth element. The investi¬ 
gation of the effect of the introduction of different 
possible fourth components on erosion resistance and 
liner performance in general would be time consum¬ 
ing. In order to keep to a minimum the number of 
firing tests, ductility as determined by mechanical 
tests should be used as the criterion of improvement. 
A high-temperature impact test, as recommended in 
Section 16.3.9, under “Notch Impact Tests and High- 
Velocity Impact Tests” would be ideal for the purpose. 
When, finally, some more ductile composition has 
been found, and its erosion resistance is to be eval¬ 
uated, care should be taken to support it adequately 
in a barrel made of high-strength steel with the bore 
surface ahead of the liner protected by chromium 


CONFIDENTIAL 




CHROMIUM-BASE ALLOYS AS LINER MATERIALS 


369 


plate or in some other suitable manner as was done 
in the tests described at the end of Section 17.4.3. 

Methods of Production 

As a parallel investigation to the continuation of 
fundamental research on the composition of chro¬ 


mium-base alloys, it is recommended that the de¬ 
velopment of suitable methods and equipment for the 
large-scale production of the alloys be continued. 

This recommendation is made in anticipation of 
the time when an alloy possessing the desired 
properties will have been developed. 


CONFIDENTIAL 





Chapter 18 

MOLYBDENUM 

By F. Palmer a 


181 INTRODUCTION 

T he search for an ideal erosion resistant material, 
discussed in Chapter 16, revealed that molybden¬ 
um was one of the few metals that showed great 
promise in this respect. In particular, it is pre-emin¬ 
ent in its resistance to thermal and chemical attack 
by the powder gases during the short time of exposure 
in a gun. At the same time, commercially available 
molybdenum lacked some of the necessary mechan¬ 
ical properties and could not be produced in the prop¬ 
er size and shape for fabricating gun liners. Extensive 
investigations have succeeded in showing ways to 
overcome these deficiencies. 

The liners which reacted most favorably to firing 
tests in a caliber .50 gun were made of an alloy of 
molybdenum with 0.1 per cent cobalt, which is harder 
than pure molybdenum. They were fabricated from 
metal which had been swaged according to a u work¬ 
ing schedule” that was developed to obtain the op¬ 
timum possible strength and ductility for a given 
reduction in cross section. These liners consisted of 
two longitudinal segments, or staves, which were 
twisted so that the seams followed the rifling grooves. 
After having been fired 2,021 rounds they were still 
serviceable, whereas gun steel would have failed after 
90 rounds under the same hypervelocity conditions 
of testing. 

Because there were still some deficiencies in the 
design of the best liners mentioned above, further 
work was planned on other methods of fabricating 
liners. Also, plans were made to prepare a molyb¬ 
denum liner for test in a 3-in. Naval gun, as described 
in Section 33.1.3. The Navy Department subsequently 
contracted with the Westinghouse Electric Corpor¬ 
ation for the use of the facilities set up under its 
contract with Division 1 and for considerable expan¬ 
sion of them according to plans already developed by 
the Division. 

182 COMMERCIAL MOLYBDENUM 

18 21 Manufacturing Process 

The term powder metallurgy is used to describe 
a The Franklin Institute, Philadelphia, Pa. 



Molybdenum 

Powder 



370 


CONFIDENTIAL 











































































COMMERCIAL MOLYBDENUM 


371 


the usual commercial process by which a molyb¬ 
denum compound, frequently the oxide (Mo0 3 ), is 
converted into bars of metal. The process starts with 
the reduction of the compound to a metal powder by 
heating it in a stream of hydrogen. The successive 
steps after that #re shown in the left-hand column of 
Figure 1. The metal powder is compressed mechani¬ 
cally in a steel mold so as to form a compact, which is 
then sintered in hydrogen at a high temperature until 
the powder particles coalesce somewhat. Further den- 
sification is accomplished by heating the sintered 
molybdenum in hydrogen to a temperature near its 
melting point (2620 C), so that the powder metal¬ 
lurgy ingot can be subsequently subjected to mechan¬ 
ical working to impart strength and ductility. The 
principal commercial forms of molybdenum are fine 
wire and thin sheet. Both are made by working such 
an ingot first hot (around 1350 C) and then gradually 
lowering the working temperature as mechanical re¬ 
duction proceeds. All heatings are made in hydrogen 
to prevent excessive loss of metal by oxidation. To 
make wire the bar is swaged in successively smaller 
dies and finally hot-drawn through other dies. To 
make sheet or strip (which is a narrow thick sheet) 
the bar is hot-rolled to the desired thickness. 

The more working molybdenum receives the more 
satisfactory become its physical properties provided 
working temperatures are properly adjusted. This 
situation is illustrated in Table 1 for sheet of different 
thicknesses made from a heat-treated compact of 
cross section lxl }/$ inches. During the working proc¬ 
ess the temperature normally falls below the point of 
recrystallization (1200 C) with the result that the 
crystal grains become elongated and interlocked, as 
shown in Figure 2. This structure in metal which is 
cold-worked below the recrystallization temperature 
promotes strength and ductility. 


18,2 2 Physical Properties of Molybdenum 

Dependence of Mechanical Properties 
on Amount of Working 

The best mechanical properties of molybdenum are 
obtained only when it is thoroughly worked. The 
increase in ultimate strength and hardness with con¬ 
tinued working, resulting in decreasing thickness, 
is shown in Table 1. The reduction in thickness of 
the sheet is roughly proportional to the amount of 
working. 


Table 1 . Dependence of physical properties of molyb¬ 
denum upon hot-working as measured by thickness of 
sheet. 


Thickness 

(in.) 

Hardness 

(D.P.H.) 

Ult. Strength 
(10 3 psi) 

Elong. (2 in.) 
(per cent) 

0.450 

227 

76.0 

slight 

0.230 

250 

88.6 

11.0 

0.015 

270 

122.0 

4.7 



Figure 2 . Structure of well-worked, commercial 
molybdenum (0.126-in. thick). Aqua regia etch, 50 X. 
(Figure 4d of NDRC Report No. A-423.) 


In the case of wire, the reduction in area is a better 
measure of amount of working. The results of some 
early tests on the ultimate tensile strength, in the 
direction of working of both wire and strip, plotted 
against cross-sectional area, are shown in Figure 3. 
Although there is a rather wide scattering of observa¬ 
tions on both sides of the mean, the general trend of 
the curve indicates definitely that the ultimate ten¬ 
sile strength increases as the metal becomes thinner 



Figure 3. Ultimate tensile strength of commercial 
molybdenum worked to different cross-sectional areas. 
(Figure 1 of NDRC Report No. A-423.) 


CONFIDENTIAL 




















372 


MOLYBDENUM 


due to the working to which it has been subjected. 
Scattering of the observations is probably due to the 
lack of uniformity in the heat-treatment of the bar 
stock. 

It may be inferred from a graph such as that in 
Figure 4 that the hardness of molybdenum also is 
related to the amount of working. In spite of the 
scatter of the observed readings, it is clear that the 
molybdenum strip became harder the more it was 
worked. 



150 200 250 300 

HARDNESS VPN 


Figure 4. Hardness of commercial molybdenum 
worked to different thicknesses. (Figure 2 of NDRC 
Report No. A-423.) 


Relation of Properties to Anisotropic 
Structure 

Although hot-working increases the strength, duc¬ 
tility, and hardness of molybdenum, the worked 
metal is likely to possess a laminated, anisotropic 
texture. The tensile strength of rolled molybdenum 
strip is approximately the same along the direction 
of working (longitudinal) as across the direction of 
working (transverse), but in the latter direction the 
ductility is very small. Hence the strip may be bent 


without cracking when the crease is transverse, but 
the metal is brittle and cracks easily when the crease 
is longitudinal. Microscopic examination of sections 
etched with aqua regia reveals a progressive change 
in crystal fabric with amount of working. Unworked 
molybdenum is fine grained with no definite crystal 



Figure 5. Structure of unworked, commercial molyb¬ 
denum (1.250-in. thick). Aqua regia etch, 50 X. (Figure 
4a of NDRC Report No. A-423.) 



Figure 6 . Structure of partially worked (laminated) 
commercial molybdenum (0.430-in. thick). Aqua regia 
etch, 50 X. (Figure 4b of NDRC Report No. A-423). 


orientation, as can be seen in Figure 5. When par¬ 
tially worked, there are crystals of various sizes which 
begin to show the effects of being worked by a ten¬ 
dency toward elongation parallel to the direction of 
working, as illustrated in Figures 6 and 7. In thor¬ 
oughly worked material the crystals are all elongated 


CONFIDENTIAL 








PRODUCTION OF DUCTILITY IN THICK SHEET LINERS 


373 


with a fibrous interlocked crystal pattern, which was 
shown before in Figure 2. 

It has been suggested 96 that circumferential duc¬ 
tility might be improved by twisting well-fibered bars 
of wrought molybdenum so that the axially elongated 
grains would be turned toward the circumferential 
direction. When suitable material and equipment be¬ 
come available, the properties of such helically fibered 
bars should be given careful study. 



Figure 7. Structure of worked but recrystallized, 
commercial molybdenum (0.274 in. thick). Aqua regia 
etch, 50 X. (Figure 4a of NDRC Report No. A-423.) 


Unsuitability of Commercial Molybdenum 
for Gun Liners 

Some of the physical properties of molybdenum 
compared with those of gun steel are given in Table 2. 
The high melting point and thermal conductivity of 
molybdenum are advantageous for its use as a bore- 


Table 2 . Physical properties of commercial molyb¬ 
denum and gun steel. 


Properties 

Molybdenum 

(wrought, 

ductile) 

Gun Steel 
(oil-quenched 
and tempered) 

Melting point (C) 

2535 

' 1450 

Thermal expansivity 

(0-500 C)(%/°C) 

5.49 X 10- 6 

13.6 X 10- 6 

Thermal conductivity 

(cal/cm/sec/°C) 

0.35 

0.102 

Specific heat (cal/g/°C) 

0.065 

0.107 

Ultimate tensile strength (psi) 

90-100 X 10 3 

>130 X 10 3 

Elongation at rupture 

(% in 2 in.) 

5-20 

>16 

Hardness at room temperature 

(VPN) 

260-270 

280-320 


surface material. The low strength and ductility of 
commercial molybdenum are undesirable properties, 
but, as will be shown in Sections 18.3 and 18.4, means 
were found for producing molybdenum and molyb¬ 
denum alloys with a great improvement in these 
properties. Even then liners had to be designed in such 
a way that the low thermal expansivity would not 
prove a major handicap, as discussed in Section 26.5.2. 

In order to withstand firing stresses of the usual 
magnitude, the metal of a gun liner must have a 
thickness of one or two tenths of an inch. On the 
other hand, sheets of commercial molybdenum that 
have been worked sufficiently to develop suitable 
mechanical properties are available only with a thick¬ 
ness of 0.015 in. or less. Hence this material is unsuit¬ 
able for gun liners. 

18 3 PRODUCTION OF DUCTILITY IN 
THICK SHEET SUITABLE FOR GUN 
LINERS 

18-31 Introduction 

The first molybdenum liners tested were tubes 
bored from swaged rods. They all failed after so few 
rounds that it became evident that molybdenum, in 
spite of its high melting point and excellent resistance 
to chemical attack by hot powder gases, would not 
prove a satisfactory liner material unless its strength 
and ductility could be largely increased. 49 - 50 This 
requirement could be met only by increasing the 
amount of working (Section 18.2.2). 

At the time that Division 1 began the development 
of molybdenum gun liners, there was no equipment 
in use that was large enough to work mechanically 
molybdenum compacts thicker than 134 in. Further¬ 
more, information was lacking as to how to work 
thicker bars without cracking them. Hence, it was 
essential to develop a working schedule which would 
change a bar of the usual dimensions (%x 134 x 20 in.) 
into a strong, ductile sheet thick enough (about 34 in-) 
for use as a caliber .50 gun liner. At the same time it 
was desirable to explore means of preparing and 
mechanically working bars or ingots of even larger 
cross section. 

18-3-2 Development of a 

Satisfactory Working Schedule 

The investigations 95 that resulted in a satisfactory 
working schedule may be summarized in terms of two 


CONFIDENTIAL 







374 


MOLYBDENUM 



Figure 8. Bend strength and yield of samples of pure 
molybdenum worked at different temperatures. (Figure 
14 of NDRC Report A-423.) 

principles. These principles are that (1) complete re¬ 
crystallization above 1200 C is essential in order to 
convert the weak, microcrystalline structure of the 
treated bars into one with large, uniform grains; and 
(2) working causes an increase in strength, by elon¬ 
gating and interlocking the large grains, only below 
1200 C , the temperature at which the metal re¬ 
crystallizes quickly. 

Development of a uniform grain structure is pro¬ 
moted by hot-rolling the bars just the proper amount 
and reheat-treating them. Since it has been found 
that crystallization is incomplete with too little work¬ 
ing and fine crystals are produced by too much work¬ 
ing, the amount of reduction in thickness on the first 
pass is critical. 

Furthermore, edge-cracking and end-splitting are 
likely to develop in the recrystallized bars unless 
they, too, are rolled once or twice above 1200 C. 
Before the temperature is reduced below this point 
it is apparently necessary to break down the equi- 
axed grain structure, which can be done only while 
the metal is plastic. The amount of reduction in 
thickness on these passes, too, is critical. 

The influence on strength and ductility of pure 
molybdenum worked under different temperature 
conditions is shown in Figure 8. The letter B indicates 
that the specimen broke at that point. The only 
specimen which did not break before the termination 
of the experiment was the one which had been worked 
at slowly decreasing temperatures. Hence, in a satis¬ 
factory working schedule provision has been made 
for the working'to take place while the temperature 
of the metal is slowly decreased. 


Table 3. Improved working schedule for alloy of 
molybdenum with 0.1% cobalt, starting with a bar 
1x114x24 in. 95 


Pass 

Furnace 

temperature 

(C) 

Reduction 
of thickness 
(in.) 

Preliminary 

1300 

0.100 

(Reheat-treat to complete recrystallization) 

1 

1300 

0.200 

2 

1300 

0.150 

3 

1200 

0.100 

4 

1200 

0.100 

5 

1100 

0.075 

6 

1100 

0.075 

7 

1000 

0.075 

8 

1000 

0.050 

9 

1000 

0.050 


Improved Working Schedule 

The best gun liners prepared for test under Divi¬ 
sion l’s auspices were produced from heat-treated 
bars of molybdenum plus 0.1% cobalt (Section 18.6), 
of dimensions about 1x134x24 in. These bars were 
processed according to the schedule shown in Table 
3. The preliminary pass served to induce crystal¬ 
lization. The structure after reheat-treatment is 
shown in Figure 9. 

The microstructure of the final product under this 
working schedule was better at a thickness of 0.250 
in. (Figure 10) than that of commercial molybdenum 



Figure 9. Structure of molybdenum alloy containing 
0 .1% cobalt, recrystallized according to improved work¬ 
ing schedule. Compare Figure 10. Aqua regia etch, 50 X. 
(Figure 13a of NDRC Report A-423.) 


CONFIDENTIAL 















HARDENING OF MOLYBDENUM BY ALLOYING 


375 


at different thicknesses (Figures 2 and 7). In addition 
it had at least 50 per cent greater strength and sev¬ 
eral times as much ductility as commercial molyb¬ 
denum of the same thickness. 

After the tests of these liners, described in Section 
18.6, evidence 76 was found of longitudinal cracking, 
spalling, and swaging of lands. This result was taken to 
indicate that even greater reduction in cross section 
will be necessary to make molybdenum a wholly 
satisfactory gun liner material, which means that the 
pressed bar must be considerably thicker than in. 

18 3 3 Advantage of Hydrostatic Compression 

The substitution of hydrostatic pressing for me¬ 
chanical pressing in the fabrication of the powder 
compact makes possible the production of thicker 
bars and also markedly improves the quality of the 
molybdenum strip. The powder, placed in a rubber 
tube, is immersed in water in a pressure cylinder and 
subjected to a hydrostatic pressure of 30,000 to 
40,000 psi. The cylinder used had a bore of 8 in., with 
which bars 3 in. square could be made. A 12-in. cylin¬ 
der was planned so that 5-in. bars could be made. 

The improvement in quality is attributable to the 
uniformity of the pressing and to the absence of 
incipient cracks. It was found that the early test 
liners made from mechanically pressed powder con¬ 
tained incipient cracks which opened up rapidly un¬ 
der firing stresses. The first crack-free liners were 
made in the form of two staves (Section 18.5) from 
molybdenum which was hydrostatically pressed. 

Later it was found possible to obtain crack-free 
material by mechanical pressing with special care 
being taken in the way the molds were filled and 
handled. It is not certain, however, that equally good 
results would be obtained from thicker bars. Further¬ 
more, mechanical pressing of large bars requires 
much more elaborate equipment than does hydro¬ 
static pressing. Hence the latter was to have been 
used in making molybdenum bars for the 3-in. gun 
liners described in Section 33.1.3. 

18 3 4 Plans for Enlarged Facilities 

It had been recognized at the inception of the 
molybdenum project that in order to fabricate mo¬ 
lybdenum for liners larger than small arms much 
more powerful equipment would be necessary than 
was then available. The decision to have such equip¬ 
ment installed at the Lamp Division of Westinghouse 


Electric Company was deferred by NDRC until 
definitive results were obtained with caliber .50 liners 
made by means of existing equipment. 

New equipment was then planned 95 for the fabri¬ 
cation of molybdenum to be used in the 3-in. gun 
liner described in Section 33.1.3. It included a 12-in. 
hydrostatic press, a heavy-duty rolling mill, and a 
7,200-ft-lb forging hammer in addition to furnaces 
and auxiliary equipment. With this equipment, b the 



Figure 10 . Structure of molybdenum alloy containing 
0.1% cobalt, reduced to 0.250 in. according to improved 
working schedule. (Compare Figure 9.) Aqua regia etch 
50 X. (Figure 13b of NDRC Report A-423.) 

sequence of operations shown on the right-hand side 
of Figure 1 could be carried out. The molybdenum 
powder used as starting material would be “doped” 
with a small amount of cobalt or other alloying con¬ 
stituent, as described in the next section. 

184 HARDENING OF MOLYBDENUM 
BY ALLOYING 

18-4,1 Hardening by Alloying Alone 

Molybdenum can be hardened not only by cold¬ 
working below the recrystallization temperature but 
also by alloying with the proper elements. The hard¬ 
ness of some alloys of molybdenum in the unworked 
condition is given in Table 4. The hardness of pure 
molybdenum in the same condition is only 175 VPN 


b The accelerated termination of Contract OEMsr-1205 
after V-J Day delayed the purchase of this equipment. It was 
later installed under a Navy Department contract with the 
Westinghouse Electric Company. 


CONFIDENTIAL 






376 


MOLYBDENUM 


Table 4. Hardness of certain alloys of molybdenum.* 


No. 

Alloy composition 

Hardness (VPN) at 
20 C 500 C 600 C 

1 

Mo + 5% W 

213 

134 

132 

2 

Mo + 10% W 

288 

160 

153 

3 

Mo + 15% W 

249 

196 

185 

4 

Mo + 20% W 

268 

191 

195 

5 

Mo + 0.01% Ni 

223 

119 

116 

6 

Mo + 7% Ni 

423 

372 

359 

7 

48.5 Mo + 48.5 W + 3.0 Ni 

494 

366 

366 


* Measurements made by Climax Molybdenum Company, 1943. 


at 20 C. These alloys have both higher ultimate 
strength and higher yield point than pure molyb¬ 
denum. Unfortunately, the last two alloys in Table 4 
are not workable alloys. 

18 4 2 Effects of Alloying and Working 

More recent observations 95 are plotted in Figure 11 
where the hardening effects of both working and al¬ 
loying with small percentages of cobalt are especially 
noteworthy. The alloy containing 0.1% cobalt is eas¬ 
ily workable and of reasonably satisfactory hardness, 
hence, in view of the difficulties involved in running 
a complete reconnaissance series, intensive work was 
spent largely in the development of this alloy. The 
improved working schedule given in Table 3 was 
devised for this particular alloy. For any other alloy 
a similar working schedule might be developed after 
a series of trials had been made. 

18-4,8 Effect of Temperature on Hardness 
of Various Alloys 

The effects of amount of working and of alloying 
with other metals are both to be seen in the curves of 
Figure 11. It will be observed that the hardness at 
20 C of unworked pure molybdenum is raised from 
about 175 VPN to about 240 VPN when worked to a 
thickness of 0.106 in.; to about 265 VPN when un¬ 
worked but alloyed with 20% tungsten; and to about 
290 VPN when alloyed with 0.1% cobalt and worked 
to a thickness of 0.25 in. The advantages possessed 
by molybdenum and its alloys over gun steel (SAE 
4150), which is harder than they are at 20 C, are two¬ 
fold: (1) above 300 C steel loses hardness rapidly 
with increase in temperature, (2) whereas with the 
other materials above 300 C the hardness falls off less 
rapidly and the rate of loss diminishes with increase 
in temperature. 

From Figure 11 it is clear that the substitution of 


an unworked molybdenum liner for one of steel will 
not eliminate swaging of the lands, for the metal is 
too soft, as was confirmed by one of the early experi¬ 
ments. 49 Improvement in this respect can be expected 
only in proportion to the degree of hardening brought 
about by more efficient working and alloying tech¬ 
niques. 



Figure 11. Hardness of alloys as a function of tem¬ 
perature—molybdenum containing different amounts 
of alloying elements compared with Stellite No. 21 and . 
SAE 4150 steel. (Figure 6 of NDRC Report A-423.) 

18 4 4 Selection of Best Alloy 

Alloys of Molybdenum with Tungsten 

The first gun liners made of a binary molybdenum 
alloy contained, respectively, 10%, 15%, and 20% 
tungsten. When tested in the caliber .50 erosion-test¬ 
ing gun (Section 11.2.1), they all failed after 150 
rounds or less. Both longitudinal and transverse 
cracks occurred, as well as severe spalling. The latter 
defect was due to their laminated structure of the 
molybdenum, which, it is now believed (Section 
18.2.2), was the result of insufficient working by an 
improved schedule. Further experiments with tung¬ 
sten as the alloying element have not been carried 
out because these alloys were found more difficult to 
work than others of nearly the same hardness, made 
at a later date, with small percentages of cobalt or 
nickel. 

Alloys of Molybdenum with Cobalt or Nickel 

Molybdenum is hardened to nearly the same extent 
by the addition of the same small amount of either 


CONFIDENTIAL 















CONSTRUCTION OF CALIBER .50 LINERS FOR TEST 


377 


cobalt or nickel, though there is evidence that nickel 
alloys harden somewhat more rapidly than cobalt 
alloys, and, therefore, are more difficult to fabricate. 

The curves of Figure 12 show how the hardness at 
20 C of such alloys, both for unworked and worked 
material, changes with the addition of increasing 
amounts of either cobalt or nickel. The change in hot¬ 
hardness of worked material at 500 C is also shown. 
Although the results shown here indicate that the 
hardening effects of cobalt and nickel are the same, 
other experiments have suggested that nickel is less 
effective than cobalt. More extensive testing is re¬ 
quired. These curves (and those in Figure 11) at least 
demonstrate that the alloy containing cobalt or nickel 
is superior to pure molybdenum, even when the latter 
is well worked. The hot-hardness is still low enough, 
however, to make the material a not entirely satis¬ 
factory one for use in a gun liner. 

Nevertheless, two caliber .50 test liners made of an 
alloy containing 0.1% cobalt gave a remarkable per¬ 
formance, as described in Section 18.6. This alloy 
does not necessarily represent the best one for the 
purpose, and therefore further study should be made 
of the hardening effects of larger percentages of cobalt 
and of varying amounts of other elements, including 
nickel, iron, and chromium. 

18 5 CONSTRUCTION OF CALIBER .50 
LINERS FOR TEST 

18,51 General Design Considerations 

The development of an improved type of molyb¬ 
denum for gun liners demanded a means of testing 
the material in an actual gun. The gun selected for 
the purpose was the caliber .50 erosion-testing gun, 
described in Section 11.2.1, because with it hyper¬ 
velocity conditions could be achieved. At the same 
time the impossibility of obtaining (with the equip¬ 
ment available) molybdenum having adequate phys¬ 
ical properties in thick sections increased the diffi¬ 
culty of making even test liners from this material. 
It was necessary to devise means of making them 
from well-worked metal of small cross section. 

Various designs for molybdenum liners were dis¬ 
cussed at the outset of the program. 52 A stave-type 
liner, that is, one having longitudinal segments, was 
suggested as a means of transferring much of the load 
of the powder pressure from the molybdenum liner 
to the steel carrier. It was hoped that, by doing this, 
molybdenum of strength and hardness less than the 


optimum might be used without severe longitudinal 
cracking during firing. Although the other designs 0 
were presumably feasible, some of them were fraught 
with more difficulties than others, and experiments 
were finally concentrated on fabricating stave liners. 



Figure 12. Hardening effect of the addition of nickel 
or cobalt to molybdenum. (Figure 7 of NDRC Report 
A-423.) 


18-5,2 Stave-Type Liners 

Stresses in Stave Liners 

A calculation of the stresses in cylindrical-tube and 
multiple-stave liners indicated that during the re¬ 
peated firing of a caliber .50 gun a cylindrical-tube 
liner would be exposed to stresses considerably higher 
than the ultimate strength of molybdenum, as is 
brought out in Section 26.5.3. Hence fatigue failure 
would be expected after relatively few stress cycles 
at the assumed explosion pressure (80,000 psi). With 
a stave-type liner, on the other hand, the stresses 
would be substantially reduced. The calculation in¬ 
dicated that the greatest reduction would occur for 
a liner nfade of four staves, although the reduction 
for a liner of two staves, in the case of the caliber .50 
gun, is nearly as great. 

Since the tangential tensile stress at the middle of 
a stave is nearly proportional to the coefficient of 
friction between molybdenum and steel, 52 it is evi- 


c The ones that were tried are mentioned in Section 18.5.3. 


CONFIDENTIAL 










378 


MOLYBDENUM 



Figure 13. Two-stave straight liner of molybdenum 
alloy containing 0.1% cobalt after having been fired 442 
rounds. (Figure 19b of NDRC Report A-423.) 


dent that a reduction in stress can be brought about 
by proper lubrication of the molybdenum-steel inter¬ 
face. Measurements 96 were made of the coefficient of 
friction between test pieces of molybdenum and SAE 
4150 steel treated with a thin film of lubricating ma¬ 
terial. Of the lubricants tested, tungsten disulphide 
was the best, with a coefficient of friction less than 
one-third that of unlubricated surfaces. It should be 
pointed out, however, that in firing a lubricated liner, 
although stresses would be reduced, any tendency 
toward forward motion and rotation of the liner would 
probably be increased. 


Seams of Stave Liners 

The first firing tests of two-stave liners 76 gave 
promise of the eventual success of this type of con¬ 
struction. The principal difficulty was that spalling, 
a severe case of which is shown in Figure 13, took 
place. Even though the staves were shrunk into the 
steel carrier with a large interference, the seams 
opened up during firing and allowed the hot powder 
gases to penetrate them and raise the temperature of 
the edges. An attempt to prevent this by brazing the 
seams was unsuccessful. Finally, marked improve¬ 
ment was brought about by hot-twisting the staves 
(1 turn in 15 in.) so that the seams ran down the 
center of the rifling grooves. 95 


Construction of Two-Stave Liners for 
Testing Materials 

After preliminary trials had showed that the best 
means of utilizing the available molybdenum sheet 
was by a two-stave liner with helical seams, a pro¬ 
gram was instituted for the trial of several series of 
such liners, each made from a different alloy. It was 
started by the test of the alloy containing 0.1% co¬ 
balt described in Section 18.4.4. The steps involved 
in construction of the two-stave liners used in those 
tests in the caliber .50 erosion-testing gun are shown 
in Figure 14. 95 The letters given in parentheses refer 
to the separate parts of this figure. 

1. Molybdenum powder containing cobalt was 
pressed mechanically and heat-treated to form a bar 
about lxlJ4 in. (Step A.) 

2. The bar was subjected to the improved working 
schedule (Table 3) to produce a strip about 0.22 to 
0.25 in. in thickness. (Step B.) 


CONFIDENTIAL 































CONSTRUCTION OF CALIBER .50 LINERS FOR TEST 


379 



Figure 14. Steps in processing a two-stave liner of molybdenum. (Figure 16 of NDRC Report A-423.) 


3. The flat strips were troughed in dies in a forging 
hammer (Figure 15) at about 1000 C. (Step C.) 

4. The troughs were tested for cracks by etching 
with hot aqua regia. Cracked troughs, an example 
of which is shown in Figure 16, were rejected. 

5. Two troughs, fitted together over a tungsten 
mandrel and held in place by a wrapping of molyb¬ 
denum wire, were twisted in a hand-twisting machine, 
illustrated in Figure 17, after having been heated to 
1000 C. (Step D.) 

6. The ends were trimmed from each twisted trough 
and the helical mating surfaces were milled as 
shown in Figure 18. 96 (Step E.) 

7. A pair of milled blanks was assembled. The out¬ 
side surface was rough-machined, and a rough bore 
was then drilled. 

8. The outside surface was ground to final dimen¬ 
sions with the assembly mounted on an arbor so that 
concentricity of bore and outside surface was at¬ 


tained. The grinding was done so that the outside of 
the assembled staves had a taper of ^ in. per foot 
and a shoulder at the breech end %-in. long with a 
1^2-in. face. Finally the ends were machined and the 
two-stave assembly was ready for insertion into the 
steel carrier. (Step F.) 

An alternative procedure for step (5) above, known 
as precision twisting, 96 consisted in mounting the 
staves, into which a close-fitting core of molybdenum 
had been slipped, in a segmented retainer which was 
then placed vertically in a steel mold and heated 
inductively to a temperature of about 550 C. The 
twisting was accomplished by hand using a cross bar 
at the top of the mold. A thin film of Aquadag baked 
on the working surfaces supplied both protection 
against oxidation and lubrication. The first staves 
which had been precision-twisted gave a performance 
inferior to those which had been hot-twisted, but 
those results are not necessarily conclusive. 


CONFIDENTIAL 










380 


MOLYBDENUM 


Deficiencies in the Design 

The foregoing construction of two-stave twisted 
molybdenum alloy liners for test purposes represents 
a first step in the utilization of this erosion-resistant 
material. Disadvantages of this design of liner are: 
(1) weakness at the seams, where spalling starts, (2) 
reliance upon the steel carrier to bear most of the 
pressure load, and (3) movement of the liner. 

Seam. The first of these deficiencies is closely re¬ 
lated to the fact that the material tested was known 
to be somewhat inadequate in physical properties, as 
explained in Section 18.4.4. When a more thoroughly 
worked molybdenum alloy becomes available, spall¬ 
ing at the seams may cease to be a difficulty. Alter¬ 
natively, the use of a seamless tube, discussed in 
Section 18.5.3, would obviate this problem entirely, 
provided that tubular liners could be made having 
the necessary tensile properties. 


Radial Load. Because the steel carrier that sur¬ 
rounds the stave liner must sustain nearly all the 
radial load from the combined gas pressure and band 
pressure (Chapter 7), the outer diameter of the liner 
is the effective inner diameter of the gun bore. Hence, 
for a stave liner to be used without reducing the factor 
of safety, there must be an increase in the wall thick¬ 
ness of the gun in the region of the liner. This change 
in turn increases the weight of the gun and reduces 
its maneuverability. 

Movement. Some caliber .50 molybdenum liners 
were observed to move both longitudinally and cir¬ 
cumferentially during firing, so that the rear joint 
opened, constriction of the bore occurred at the for¬ 
ward end, and the rifling of the liner no longer regis¬ 
tered with that of the barrel ahead of it. This sort of 
movement, which was due to insufficient anchorage 
of the liner in its steel carrier, was corrected by add¬ 
ing a shoulder and increasing the interference of the 



Figure 15. Troughing operation in making a two-stave liner of molybdenum. (Figure 17 of NDRC Report A-423.) 


CONFIDENTIAL 





CONSTRUCTION OF CALIBER .50 LINERS FOR TEST 


381 


shrink-fit. Nevertheless, even the best liners still 
showed a tendency to forward movement of the liner 
with consequent constriction of the bore at the for¬ 
ward end. It was not possible to determine from the 
tests conducted whether this fault would also be elim¬ 
inated by the use of material of greater strength, 
although there is a strong a priori assumption that it 
would be. 

18 5 3 Other Designs of Molybdenum Liners 95 

As has already been mentioned, the design of a gun 
liner to be made from the molybdenum thus far avail¬ 
able has had to take into account the lack of strength 
of the material and its anisotropic character. Al¬ 
though liner design in general is discussed in Chapter 
26, it is appropriate to record here the experiments 
that were made in an effort to develop a practical 
means of testing molybdenum in the caliber .50 ero¬ 
sion-testing gun and to explore the potentialities of 
this new gun liner material. 

Brazed-in One-seam, Tube of Molybdenum Sheet. A 
test piece was prepared by brazing a lining of 0.025- 
in. thick molybdenum into an SAE4150 steel carrier 
with a special copper-base brazing alloy in an atmos¬ 
phere of Ammogas (NH 3 ) and heat-treating it. Hy¬ 
drostatic pressure was applied by filling the bore with 
Wood’s metal held in place by close fitting steel plugs 
at the ends. Fracture of the lining began after ap¬ 
proximately 200 stress cycles, and increased to the end 
of the test (1,020 cycles). A later circumferential bend¬ 
ing test revealed the transverse weakness so char¬ 
acteristic of molybdenum. This method of fabrica¬ 
tion merits further study with hardened molybdenum- 
alloy sheet. 

Tube Made from Flat Disks. The disk construction 
was designed to minimize the radial distortion of the 
carrier assembly due to firing pressure by having the 
molybdenum oriented in such a way that its direction 
of greatest strength opposed the radial stresses. A 
thick-walled tube was made by brazing a stack of 
punched disks, machining the outside for a press-fit 
into the carrier and machining the inside to receive a 
tapered tubular caliber .50 liner made from a bar of 
swaged, recrystallized, pure molybdenum. The braz¬ 
ing material, prepared especially for this purpose, was 
a copper-silver alloy fortified with palladium. The 
tubular liner, with an interference of 0.003 in., stuck 
during insertion and defied all attempts to remove it 
intact. Recovered fragments showed indentations by 
the edges of the disks caused by the shrink-fit stresses, 


which would have keyed the liner in position. Another 
liner of this type should be made for test firing. 

Wire-Wrapped Tube. Another method of orienting 
molybdenum in such a way that the directional char¬ 
acteristic of its strength is made use of consists in 
wrapping molybdenum wire around a tube made from 
molybdenum sheet and brazing the wire in place. 
Some molybdenum wire having a trapezoidal cross 
section was drawn for this purpose by means of a 
Turk’s head machine, but a liner was not actually 
constructed. This design also has the drawback that 
some of its strength would depend on the brazing 



Figure 16. Cracks in a molybdenum trough disclosed 
by hot aqua regia etch. About 2X. (Figure 18 of NDRC 
Report A-423.) 


compound, which might lose strength at gun-oper¬ 
ating temperatures. 

Cast-Bonded Molybdenum-Steel Strip. Bimetallic 
slabs were made by casting SAE4150 steel against 
one face of a flat strip of hot-rolled molybdenum 
fluxed by painting with a suspension in acetone of 
finely ground, dehydrated, Easy-Flo flux. Pieces for 
tension and fatigue tests were made from these com¬ 
posite slabs hot-rolled to about a quarter of their orig¬ 
inal thickness (final thickness: Mo, 0.015 in.; steel, 
0.125 in.). The results of these tests were encouraging 
so far as resistance to stress was concerned, but the 
interfacial bond, although apparently complete, 
showed under the microscope nodules of iron-molyb¬ 
denum alloy which extended into the molybdenum 
and acted as stress-raisers. There was also a thin, 


CONFIDENTIAL 





382 


MOLYBDENUM 



Figure 17. Twisting operation in making a two-stave liner of molybdenum. (This figure had accompanied the manu¬ 
script of NDRC Report A-423, but it was not reproduced in that report.) 


brittle layer between steel and molybdenum that was 
a definite source of weakness. d 

Seamless Tube. An experiment was conducted in an 
effort to evaluate the possibility of making a seam¬ 
less gun liner of molybdenum. It involved extrusion 
of hot molybdenum through a die. Although a seam¬ 
less tube could not actually be made with the equip¬ 
ment available, it was possible to begin extrusion of 
molybdenum at a temperature of 1200 C. This exper¬ 
iment indicated that pure molybdenum would have 
to be maintained at that temperature during the ex¬ 
trusion process, which would make it extremely dif¬ 
ficult to find suitable dies. 

A more promising type of tube-making process is 
one involving hot-rolling. The plans for continuation 
of the molybdenum project by the Westinghouse 

d Perhaps this layer may be the same as that encountered by 
other observers and described in Sections 18.8.2 and 21.3.3. 


Company under Navy contract included the in¬ 
stallation of tube-making equipment. 

18 6 FIRING TESTS OF CALIBER .50 
MOLYBDENUM LINERS 

18 61 Early Liners 

The first molybdenum liners tested were fired in a 
caliber .50 machine gun at the Geophysical Labora¬ 
tory (Section 11.2.2). 49 - 50 They were bored from 
swaged molybdenum rods. Although longitudinal 
cracking and swaging of the lands caused failure after 
relatively few rounds, the tests served to demonstrate 
the remarkable inertness of molybdenum with re¬ 
spect to the powder gases. 

Subsequent testing of molybdenum liners was car¬ 
ried out in the caliber .50 erosion-testing gun (Section 


CONFIDENTIAL 






FIRING TESTS OF CALIBER .50 MOLYBDENUM LINERS 


383 



11.2.1) at the Franklin Institute. 76 The first 24 liners 
tested failed for mechanical reasons, such as cracking 
or constriction of the bore, simply because the molyb¬ 
denum did not have sufficient strength and suffi¬ 
ciently uniform metallographic structure. The tests did 
confirm the resistance of molybdenum to erosion by 
powder gases, even those from powder containing 
40% nitroglycerin. 

Most of this group of liners were of the stave type 
(Section 18.5.2), including two with four staves and 
two with ten staves. Although the unsuitability of 
the material prevented the tests from being definitive, 
they did furnish clues to improvements in design of 
test liners, which were incorporated in the design 
described in Section 18.5.2. 

The material used in 13 of these liners was pure 
molybdenum. Five of them were made from alloys 
containing 5 to 20% of tungsten. The remainder con¬ 
tained a few hundredths of 1% of nickel as alloying 
constituent. Because none of this material had been 
fabricated by the improved working schedule de¬ 
scribed in Table 3, the possible value of these alloying 
constituents could not be assayed. After the improved 
working schedule had been developed, it was planned 
to make a systematic study of the effect of different 
alloying constituents. 


18-6,2 Liners of Molybdenum 

Hardened with Cobalt 

From the first batch of molybdenum alloy hard¬ 
ened with 0.1% cobalt and fabricated according to 
the improved working schedule (see Section 18.3.2) 
four two-stave liners were constructed. 96 The physi¬ 
cal properties of the material in these liners is given 
in Table 5. The measurements were made on test 


Table 5. Physical properties of molybdenum hardened 
with cobalt used for caliber .50 test liners. 95 


Liner No. 

Strength* 

(%) 

Yield f 

(%) 

Hardness J 

BL 33-3 

138 

.95 

241 

BL 33-4 

150 

1.66 

235 

BL 33-5 § 

166 

3.5 

226 

BL 33-7 § 

148 

1.08 

234 


* The ultimate strength compared with that of factory-worked molyb¬ 
denum of the same gauge thickness as determined by a simple bend test, 
t The yield at rupture as determined in the same test, 
t Vickers pyramid number at room temperature. 

§ Helical seams. 

specimens cut from the ends of the strips from which 
the staves were constructed. 

Two of the liners had straight seams and two had 
helical seams. All had a j^-in. shoulder, an average 


Figure 18. Milling the mating surfaces of one stave of a two-stave molybdenum liner. The stave was held in a fixture 
that was rotated at the same time that it was moved forward. (Figure 73 of NDRC Report A-424.) 


CONFIDENTIAL 






















384 


MOLYBDENUM 



Figure 19. Two-stave twisted liner of molybdenum 
alloy fired 2,024 rounds. See Figure 4 of Chapter 16. 
(Figure 19a of NDRC Report A-423.) 


wall thickness of 0.13 in., and a diametral taper of 
34 in- Per ft. 

These four liners were tested in the caliber .50 
erosion-testing gun. Ball bullets were fired at a muz¬ 
zle velocity of about 3,700 fps, using a double-base 
powder (20% nitroglycerin) at a maximum pressure 
of about 57,000 psi. 

The results of the tests are summarized in Table 6. 
They represented a tremendous improvement, for the 
longest previous test had lasted only 454 rounds. The 
fact that liner No. BL 33-4 lasted only 152 rounds, 
however, indicated that the molybdenum was still 
not as uniform as desired. 


Table 6. Firing tests in caliber .50 erosion-testing gun 
of two-stave liners of molybdenum hardened with 0.1% 
cobalt. 76 95 


Liner No. 

Rounds 

Cause of failure 

BL 33-3 

1133 

Bore constriction 

BL 33-4 

152 

Spalling 

BL 33-5 

2022 

No failure 

BL 33-7 

2024 

No failure 


The tests of liners BL33-5 and BL33-7 were 
arbitrarily stopped, although the liners were in 
condition to withstand further firing. There was 
no evidence of powder gas erosion of the bore 
surface. 

Tests of five other liners made from different 
batches of molybdenum hardened with 0.1% cobalt 
were not carried as far as those of liners BL 33-5 and 
BL 33-7. Some of them were made to test special 
design features, such as the effect of precision twist¬ 
ing (Section 18.5.2). The overall result of the nine 
tests of hardened molybdenum liners was that the 
following types of failure were still present, although 
to a much smaller degree than with the first group of 
liners. 

Cracking. Longitudinal cracking usually occurred 
in the center of the grooves after a relatively few 
rounds. The lack of sufficient transverse strength and 
ductility permitted the grains, elongated by working, 
to split apart. This effect was probably aggravated 
by the smallness of the coefficient of expansion of 
molybdenum compared with that of steel, for the 
transverse cracking that took place across the axis of 
the grain structure would not have occurred if the 
liner had received proper support from the assembly 
at all temperatures. Surface checkerwork cracking 
(seen in Figure 19) was caused by the thermal stresses 
at the bore surface. 


CONFIDENTIAL 






















OTHER METHODS OF FABRICATING MOLYBDENUM 


385 


Spalling. Spalling occurred along the edges of seams, 
particularly where the lands crossed straight seams. 
It often emanated from tool marks and incipient 
cracks in the surface (Figure 16) which are the result 
of cleavage fractures in individual crystals under the 
stresses imposed by the cutting tool. Preliminary ex¬ 
periments 96 have demonstrated the feasibility of re¬ 
ducing both amplitude and asperity of such imper¬ 
fections by electropolishing, but the problems in¬ 
volved in electropolishing the bore of liners after in¬ 
sertion into the carriers have yet to be solved. 

Swaging of Lands. Swaging of the lands increased 
the land diameter at the origin of rifling and for a short 
distance beyond. In the case of liners BL 33-5 and 
BL 33-7 the land height at the origin of rifling was re¬ 
duced from 0.010 in. to approximately 0.007 in. at the 
end of the test. This increase in diameter produced a 
drop in both pressure and velocity. The cause of 
swaging was a lack of sufficient hot-hardness to with¬ 
stand engraving stresses. Swaging, however, was re¬ 
duced to a remarkable degree by the use of the alloy 
containing 0.1% cobalt. 

Liner Movement. Forward movement of the liner 
opened the joint at the breech end beneath the cart¬ 
ridge case and extrusion of the brass case into this 
opening sometimes prevented extraction of the case 
and stopped the test. All liners made without an 
integral shoulder (Figure 14F) and fired a long 
time moved forward and opened up at the rear 
joint. 

Plastic flow of metal toward the muzzle end of the 
liner caused bore constriction which on some occa¬ 
sions terminated the test. In this case, too, the in¬ 
creased hardness of the improved alloy decreased this 
type of failure. 

All these types of failure were influenced if not 
caused entirely by the fact that the physical proper¬ 
ties of the molybdenum were not optimum. The ex¬ 
tent to which the design of the liner itself may have 
contributed has already been considered in Section 
18.5.2. These tests, then, represent merely a begin¬ 
ning in the development of molybdenum as a suitable 
material for gun liners. The program planned by 
Division 1 and subsequently taken over by the Navy 
Department envisaged extensive tests of caliber .50 
liners for the double purpose of determining whether 
changes in the fabrication process might give im¬ 
proved material and of trying preliminary design fea¬ 
tures that might be incorporated in a molybdenum 
liner for a medium caliber gun, such as that described 
in Section 33.1.3. 


18 7 OTHER METHODS OF FABRICATING 
LARGE BILLETS OF MOLYBDENUM 

18,7-1 Introduction 

Some of the foregoing sections of this chapter have 
told of the need for large presses and furnaces in order 
to fabricate large billets of molybdenum by the usual 
powder metallurgy process. Therefore alternative 
methods were explored by Division 1 for fabricating 
such billets as starting material for making gun liners. 
In three of them molybdenum was melted; in the 
fourth it was merely sintered. The products resulting 
from the experiments carried out were too brittle for 
the desired purpose; but there is still a possibility 
that either vacuum melting or inductive sintering in 
an atmosphere of hydrogen might be developed to a 
successful conclusion. 

18,7,2 Vacuum Melting in an Electric Arc e 

In this method of making a large billet of molyb¬ 
denum an arc is established in vacuo between vertical 
electrodes made from compressed molybdenum pow¬ 
der. The molybdenum runs down from the upper 
electrode and collects at the bottom of the furnace in a 
casting mold that is lined with molybdenum sheet. 
Although the process involves the liberation of energy 
at high concentration, it is readily controlled, makes 
the use of refractories unnecessary, and yields a 
nearly pure product. 

Ingots as large as 25 lb have been cast. Ingots of 
either molybdenum or molybdenum-base alloys are 
found to contain fewer nonmetallic inclusions than 
metal processed in other ways, but they contain 
many gas pockets, caused by the escape of water 
vapor, hydrogen, and carbon monoxide which orig¬ 
inated in the bar stock. By remelting in a vacuum-arc 
furnace such an ingot was densified so that the final 
product had a density of 10.232 g/cu cm, which is 
nearly that of the maximum theoretical density of 


e This process was being developed by the Climax Molyb¬ 
denum Company at its Research Laboratory in Detroit at the 
time that Division 1, NDRC, entered into Contract OEMsr- 
1273 for the development of chromium-base alloys (see 
Chapter 17). The scope of that contract was made to include a 
study of the vacuum melting of molybdenum in order to de¬ 
termine whether the Climax Company’s process might be 
applicable to the preparation of molybdenum for gun liners. 
The results of the investigation were described in a series of 
monthly progress reports. They were summarized in a formal 
Division 1 report 97 that was drafted but not issued. The Climax 
Company planned to continue the investigation for its own 
purposes after the termination of Contract OEMsr-1273. 


CONFIDENTIAL 




386 


MOLYBDENUM 


molybdenum (10.295 g/cu cm), computed from x-ray 
lattice parameter measurements. 

The microstructure of one of these pure molyb¬ 
denum castings, shown in Figure 20, consists of large, 
columnar grains similar to those in the recrystallized 
material made by powder metallurgy technique, 
shown in Figure 9. The grains of the molybdenum al¬ 
loy castings are finer and equiaxed, but in both cases 
the grain size is larger than that of the original com¬ 
pressed and sintered bars. Hence it is not surprising 
that when an attempt was made to forge this mate¬ 
rial the ingot split into many rodlike pieces. 95 



Figure 20. Structure of molybdenum remelted by the 
vacuum-arc process. 50 X. (This figure had appeared 
in a progress report on Contract OEMsr-1273.) 


Molybdenum-base alloys containing tungsten, iron, 
or cobalt when melted in the vacuum arc are not as 
hard as similar alloys prepared by powder metal¬ 
lurgy. 

18 7 3 Vacuum Melting 

in an Induction Furnace 96 

Melting Molybdenum Alloys in a Crucible of 
Pure Aluminum Oxide 

A vacuum furnace heated by induction has been 
used successfully in melting molybdenum-base alloys 
containing chromium, cobalt, and iron. The details 
of the furnace may be seen in Figure 21 which is here 


shown as used in the melting of molybdenum-base 
alloys with a molybdenum content of 50 to 80%. The 
Alfrax crucibles, consisting of pure aluminum oxide 
(A1 2 0 3 ), melt at a temperature in the neighborhood of 
2000 C, hence they cannot be used to melt pure mo¬ 
lybdenum (melting point, 2620 C). Energy was sup¬ 
plied to the induction coils from a vacuum-tube oscil¬ 
lator generator rated at 20 kw and running at 
approximately 130 kc. 



Figure 21. Vacuum-melting furnace for molybdenum 
alloys. (Figure 40 of NDRC Report A-424.) 


The compacts (lj^ in. x 2 to 3 in.) were melted in 
about 6 min and allowed to cool in the crucibles. 
When the melts were removed from the furnace they 
showed bright untarnished surfaces, but out of about 
50 alloys with varying compositions, all but a dozen 
were too brittle to be of use. The hardness, strength, 
and malleability of the better ones was not definitely 
superior to those of molybdenum except in the case 
of one alloy which, upon analysis, had the following 


CONFIDENTIAL 




























































OTHER METHODS OF FABRICATING MOLYBDENUM 


387 


composition: Mo 79.4%; Cr 8.5%; Co 12.7%. This 
material, or some similar alloy, might be useful as a 
gun liner material when adequately supported by the 
gun tube proper. 

Melting Molybdenum in Thorium Oxide 
Refractories 

Test samples of molybdenum alloys have been 
melted successfully in an inductively heated vacuum 
furnace having an enclosure of silica glass surrounded 
by high-temperature thermal insulation of calcined 
thorium oxide. This hearth material is little affected 
by temperatures considerably above the melting 
point of molybdenum. In one test, melting was ef¬ 
fected with a 1,100-g charge in a crucible 2%; in. in 
internal diameter in the field of a coil of 6-in. internal 
diameter by 10 in. long operated at 800 to 1,000 
ampere-turns per inch at a frequency of 9,600 c. In 
another test, melting was brought about progres¬ 
sively, using the same power supply, in a charge ap¬ 
proximately 134 in. in diameter and 12 in. long 
formed by stringing drilled blocks of sintered molyb¬ 
denum on a molybdenum rod to maintain the align¬ 
ment necessary for subsidence of the column as melt¬ 
ing progressed. 

Most of the experimental ingots produced by this 
melting technique were brittle and showed extensive 
intercrystalline cracking when forged at approximate¬ 
ly 1400 C. This may have been due to an observed 
nonmetallic material of purplish hue and unknown 
composition 1 which appears to form an extensive 
intercrystalline network. The crystalline structure of 
the ingots was coarse and columnar. In one case 
single crystals about 1 cm in length were detached, 
and at room temperature were bent repeatedly 
through an angle of 30 degrees provided the rate of 
deformation was slow. However, a suddenly applied 
bending force always caused fracture. 

A mixture of molybdenum with thorium oxide 
(Th0 2 ), mechanically pressed and sintered at 1000 C, 
was melted in similar manner. It was thought that 
the thorium oxide might minimize the grain size in 
the ingot. The grains were found to be equiaxed, but 
pronounced grain boundary attack resulted from 
etching in aqua regia. 


f This is, perhaps, the same material described in a progress 
report dated March 27, 1944, by the Climax Molybdenum 
Company as follows: “. . . a purple tinted material having 
characteristics between metal and oxide. Its analysis . . . gives 
it almost exactly the composition of MoO.” 


Thermit Meltings 

For some time commercial ferro-molybdenum has 
been produced by “thermit” melting. In attempting 
to adapt this technique to the production of ingots of 
molybdenum of high purity all ferrosilicon was dis¬ 
pensed with and aluminum alone was used to reduce 
the molybdenum oxide (M0O3) to the metallic state. 
Carefully calculated amounts of molybdenum oxide 
and aluminum were placed in a pot over a form lined 
with periclase (MgO) and surrounded by a steel shell. 
After having been baked out both form and shell 
were buried in a sand pit with the top flush with the 
sand level. 

The charge was ignited with a bomb containing 
about 34 lb of aluminum and about 2 oz of Solazone 
(Na 2 0 2 ) which could be lighted with a taper. The 
reaction began at once and proceeded rapidly, only 
from 10 to 30 sec having been required for the com¬ 
bustion of about 200 lb of mixture. During the main 
reaction large volumes of smoke and flames were 
produced, as shown in Figure 22. This stage was fol¬ 
lowed by a second, of longer duration, during which 
the melt bubbled and emitted only a small amount of 
smoke and flame. After cooling for 24 hr the ingot 
was removed and, if possible, sawed in two. If it was 
too hard to saw, it was broken with a sledge hammer. 

The characteristics of such ingots are shown in 
Figure 23 and of the accompanying slags in Figure 
24. Mixes Nos. 35, 36, and 38 contained a small 
amount of iron or iron ore as an oxidizing agent, but 
No. 37 contained no iron. All four ingots were de¬ 
scribed as very porous, soft, gassy, and coarse grained. 
The top of No. 37 was fine grained and contaminated 
with hard slag. Segregation is seen to be more pro¬ 
nounced in No. 38. Such material is evidently unsuit¬ 
able for use as a gun liner. 

In the experiments performed it was found impos¬ 
sible to obtain a balanced mix which would give a 
degassed product without an excess of silicon or alum¬ 
inum or both, elements which tend to make the alloy 
hard and brittle. Equally unsuccessful was the at¬ 
tempt to find some other material (iron ore, iron, 
chromium, manganese, copper) which, when added 
to the mix, would perform the final deoxidation and 


g This process was investigated by the Climax Molybdenum 
Company at its plant at Llangeloth, Pennsylvania. The results 
were described in three reports (dated March 27, July 1, and 
September 1, 1944) submitted to Division 1, NDRC, under 
Contract OEMsr-1273. They were summarized in a formal 
Division 1 report 97 that was drafted but not issued. 


CONFIDENTIAL 





388 


MOLYBDENUM 


leave in the melt a residual element which would not 
affect its properties adversely. 



Figure 22. Thermit process in which 100-lb ingot of 
molybdenum is being prepared by reduction of technical 
molybdenum oxide with aluminum alone. The picture 
was taken at 1/2000 sec at/ll about 5 sec after start of 
reaction. (This figure had appeared in a progress report 
on Contract OEMsr-1273.) 

18 7 5 Inductive Sintering in an 

Atmosphere of Hydrogen 96 

Inductive heating appears to be a practicable 
means for heating molybdenum or molybdenum alloy 
(possibly also tungsten) compacts of large cross sec¬ 
tion to put them into condition for subsequent work¬ 
ing. In certain exploratory experiments a 13^x3J4-in. 
compact was formed by pressing molybdenum pow¬ 
der mixed with 0.1% of Steartex, for lubrication and 
increased coherence, in a cylindrical die at 30 tons per 
square inch. The compact was embedded in 40-mesh 
white fused alumina contained in the quartz tube 
furnace shown in Figure 21, in which a hydrogen at¬ 
mosphere was maintained. Temperatures were de¬ 
termined by viewing a cavity in the top of the com¬ 


pact through a sight tube by means of an optical 
pyrometer. A current of approximately 1,400 ampere- 
turns per inch at a frequency of 9,600 c was required 
to raise the temperature of the compact to approxi¬ 
mately 2000 C where it was held for about 5 min. The 
total heating period was about 90 min. Better sinter¬ 
ing of the interior of the billets would be obtained by 
the use of the maximum practicable sintering tem¬ 
perature and an increase in the time the billets are 
held at such temperature. 



Figure 23. Sawed and split ingots of molybdenum 
melted by the thermit process. See Figure 24 for the slag 
from the same heat. (This figure had appeared in a 
progress report on Contract OEMsr-1273.) 



35 36 37 38 


Figure 24. Specimens of slag from same four heats 
from which specimens of molybdenum in Figure 23 were 
obtained. (This figure had appeared in a progress report 
on Contract OEMsr-1273.) 

After removal from the furnace measurements of 
the cold sintered billet showed that it had undergone 
a volume shrinkage of about 40 per cent, and had a 
density of approximately 9.5 gm/cm 3 . It was forge¬ 
able, and when worked at temperatures from 900 to 
1100 C yielded a sound machinable bar Y^xY^xli in. 
Such bars were found to be likely to show end¬ 
cracking but no other defects. A test of one of the fin¬ 
ished pieces at room temperature gave a hardness of 
259 YPN. Unlike most metals and alloys, the tensile 


CONFIDENTIAL 









COATINGS OF MOLYBDENUM 


389 


properties of molybdenum, at room temperature, 
appear to be highly responsive to variations in the 
rate of deformation, the higher values of ultimate 
strength and lower values of ductility accompanying 
the more rapidly applied stress. 

18 8 COATINGS OF MOLYBDENUM 
1881 Introduction 

Instead of applying molybdenum to the bore sur¬ 
face of a gun in the form of a liner, consideration has 



Figure 25. Structure of molybdenum coat sprayed on 
steel with wire gun. The crystals were flattened by spat¬ 
tering. Etched with KOH-i- K 3 Fe(CN) 6 ,250 X. (Figure 3 
of NDRC Report A-418.) 

been given to the possibility of applying a coating on 
steel or other base metal. Unsuccessful attempts to 
obtain satisfactory molybdenum coatings by electro¬ 
deposition of metal from aqueous solutions are men¬ 
tioned in Sections 16.4.4 and 20.1.3. No satisfactory 
deposits were obtained by aqueous plating of suboxide 
and subsequent reduction in hydrogen. 

The exploratory attempts at spray coating are 
described in the next section. The more extensive ex¬ 
periments with deposition of a coating of molybdenum 
by pyrolytic plating are described in Chapter 21. 

18 8 2 Spray-Coating Molybdenum 90 

Molybdenum coats of different thicknesses can be 
sprayed on roughened steel, ceramic, graphite, or pa¬ 
per mandrels with a “gun” using wire and a 


mixture of oxygen and acetylene under pressure. 
Such a coating on steel is porous and has a flakelike 
structure with flattened grains due to plastic defor¬ 
mation during deposition, as shown in Figure 25. 
Sintering in hydrogen produces only partial densifi- 
cation and the resulting coat is very brittle. Densi- 
fication by fusion with an atomic hydrogen arc gives 
rise to coarse-grained coats more brittle still, due to 
the formation of an intermetallic compound at the 
interface. 11 Complete densification may be accom¬ 
plished by spraying on top of the molybdenum a thin 
top-coat of nickel, or nickel and copper, or stellite, 
and sintering in hydrogen above the melting point of 
the top-coat. The coatings so formed have very little 
ductility. 

Densification and improved bonding of the molyb¬ 
denum to the underlying steel can be accomplished 



Figure 26. Structure of sprayed coat of molybdenum 
roller-welded to steel. Note heat effects to a depth of 
0.1 in. below coat. Nital etch, 200 X. (Figure 4 of NDRC 
Report A-418.) 

by the use of an electric resistance roller welder. The 
outside of the steel liner rests upon one roller, while a 
second roller is pressed against the molybdenum coat¬ 
ing inside. When the liner is revolved and a current 
of several hundred amperes is passed through coat 
and liner from one roller to the other, the coat is 
heated and welded to the steel. Figure 26 illustrates 
the microstructure of a welded coat. No roller-weld 
coat was made with properties sufficiently good to 
justify a firing test. 


h See footnotes (d) and (f). 


CONFIDENTIAL 







390 


MOLYBDENUM 


Another method of making a coat of molybdenum 
on steel is by evaporation from an atomic hydrogen 
arc, or by projecting the powder into such an arc. The 
very high temperatures involved in such experiments 


bring about the formation of intermediate layers of a 
brittle intermetallic compound which destroys the 
usefulness of the coat as an erosion-resistant gun 
lining. 


CONFIDENTIAL 



Chapter 19 

STELLITES AND OTHER COBALT ALLOYS 

By J. F. Schairer a 


191 INTRODUCTION 

T he survey of erosion-resistant materials given in 
Chapter 16 mentions that Stellite No. 21, because 
of its resistance to chemical attack by hot powder 
gases, hardness and strength at high temperatures, 
and excellent wear and abrasion resistance, has shown 
outstanding performance in applications as short 
breech liners in machine gun barrels under firing con¬ 
ditions so severe that unmodified steel barrels are un¬ 
able to withstand such schedules. The present chap¬ 
ter describes the metallurgy of Stellite No. 21 and 
then recounts the tests that have been made with 
other stellites and other cobalt alloys in an effort to 
utilize them as gun liner materials. The conclusion 15 
from those investigations was that “Stellite No. 21 
is an alloy composition that happens to have a nearly 
unique combination of properties that makes it suit¬ 
able for use as a gun liner fired with single-base 
powder.” 

At the same time, the advantages and limitations 
of Stellite No. 21 as a liner material should be clearly 
recognized. Among the advantages are its hot-hard- 
ness and hot-strength (to resist deformation of the 
rifling), its excellent abrasion and wear resistance (to 
resist wear by the projectile), its good ductility 
(which prevents serious cracking) and its machinabil- 
ity (which permits satisfactory fabrication). The low 
thermal conductivity of Stellite No. 21 as compared 
to gun steel (about that of gun steel at 200 C and 
about % at 600 C) may be an advantage in some 
cases for a gun with a stellite liner where the temper¬ 
ature of the steel gun tube limits performance because 
of lack of strength at high temperatures. 

The low melting points (actually fusion ranges) of 
stellite-type alloys are an outstanding limitation. The 
melting points of these alloys lie in the range between 
1250 and 1350 C while gun steels are in the range of 
1400 to 1450 C. There is some evidence (described in 


a Special Assistant, Division 1, NDRC. (Present address: 
Geophysical Laboratory, Carnegie Institution of Washington.) 

b This conclusion was expressed in the “Report of Stellite 
Advisory Committee” of Division 1, NDRC, October 5, 1945, 
after detailed study of the results of the investigations that are 
summarized in the present chapter. 


Chapters 10 and 13) that erosion, particularly in 
large guns, is, in part, an actual melting of the bore 
surface during firing. In all tests of stellite liners in 
the caliber .50 erosion-testing gun, in machine guns, 
and in the 37-mm gun, M3, when double-base powder 
(with a high adiabatic flame temperature) was used, 
serious melting of the bore surface occurred; while in 
similar tests using a cooler propellant (IMR or FNH- 
M1 powder) performance was excellent. Because of 
their low melting points stellites are not a univer¬ 
sal solution to the problem of the erosion of a 
hypervelocity gun. 

19 2 SELECTION OF STELLITE NO. 21 
AS A LINER MATERIAL 

19,2-1 Vent-Plug Tests of Stellite No. 6 

At a very early stage of the studies of gun erosion 
by Division 1,° NDRC, the suggestion was made by 
C. W. Drury, Chief, Armour Plate Division, Depart¬ 
ment of Munitions and Supply, Canada, that gun 
erosion might be mitigated by use of bore-surface 
material having high hot-hardness. He suggested 
further that cobalt alloys (for example, the stellites) 
were the materials to investigate, since cobalt is the 
element that confers this property on an alloy. 

As a result of this suggestion, and in spite of criti¬ 
cism of it expressed at a conference 297 with ordnance 
experts, both Stellite No. 6 and high-speed steel 
(which also has a certain degree of high hot-hardness) 
were tested as erosion vent plugs (Section 16.3.1). 
Both materials showed high weight losses. 27 The 
evident need for a material with a high melting point 
to withstand this test was partly the reason that 
attention was focused on the development of chro¬ 
mium (Chapter 17) and molybdenum (Chapter 18). 
It was not until later that it was realized that the 
poor resistance of stellite to “flash” melting in this 
very severe vent-plug test had masked its resistance 
to chemical attack by the powder gases, a conclusion 
that was confirmed by subsequent vent-plug tests of 
several stellites under less severe conditions. 75 


c At that time Section A, Division A of NDRC. 


CONFIDENTIAL 


391 




392 


STELLITES AND OTHER COBALT ALLOYS 


19,2,2 Machine Gun Liner Tests of 
Stellite No. 6 

A firing test on a short breech liner of pure soft 
molybdenum in the caliber .50 air-cooled, heavy ma¬ 
chine gun barrel showed that, even though there was 
no melting or chemical attack, failure occurred by 
deformation of the rifling by the swaging impact of 
the bullets after only a short burst of fire. These test 
results emphasized the importance of hot-hardness, 
resistance to permanent deformation, and wear re¬ 
sistance of the bore-surface material in rapid-fire 
guns even when fired under ordinary velocity condi¬ 
tions. 

In order to evaluate the relative importance of hot¬ 
hardness and chemical resistivity in controlling the 
performance of bore-surface materials in a rapid-fire 
gun, a reconnaissance series of different materials of 
known hot-hardness was selected for firing tests as 
short breech liners in the caliber .50 heavy machine 
gun barrel. Liners of 18-4-1 high speed steel, Voland 
No. 2 hot-die steel, Cyclops KL hot-die steel, d and 
Stellite No. 6 were tested in the caliber .50 heavy 
machine gun barrel on a severe firing schedule (500 
round groups of 100-round bursts one minute apart). 
The results 80 on the first three materials were very 
disappointing. The hot-hardness of these materials 
could not be utilized because of severe powder gas 
erosion. Evidence for the thermal alteration at the 
bore surface of one of the liners of hot-die steel is 
shown in Figure 9 of Chapter 16. The liner of Stellite 
No. 6, on the other hand, was resistant both to therm¬ 
ochemical attack by the powder gases and to de¬ 
formation of the rifling. 

Since a test of materials of outstanding hot-hard- 
ness was desired, the particular stellite alloy chosen 
for the test (Stellite No. 6), had been selected because 
it was the hardest commercial grade of stellite that 
could be rifled with carbide-tipped cutting tools and 
because it had an excellent record as a valve-seating 
material for high-temperature service. After three 
500-round groups, the stellite liner showed little or 
no powder-gas erosion, deformation, or wear. A gun 
steel barrel in a similar test is nearly worn out after 
only one such group. However, the liner of Stellite 
No. 6 showed several deep longitudinal cracks, at 
least one of which passed completely through the 
liner wall. Such a liner would be unsafe for Service 
use. Hot-hardness had been overemphasized at the 
expense of ductility; moreover, since the liner had 

d For the compositions of these steels see Section 16.4.11. 


moved forward during firing, the insertion problem 
remained to be solved. 

19 2 3 Tests of Stellite No. 21 

The insertion problem was solved by the use of the 
flanged liner design described in Section 22.2.1. The 
problem of cracking was completely solved by chang¬ 
ing the composition of the stellite for liner use. Using 
the same severe firing schedule, a 6-in. length invest¬ 
ment-cast liner of Stellite No. 21 was fired 10,900 
rounds before it finally wore out. e There was no crack¬ 
ing or dangerous failure. This alloy had good hot¬ 
hardness, but not as great as that of Stellite No. 6. 
The ductility of Stellite No. 21, however, is so much 
greater than that of No. 6 that surface cracks are not 
propagated in it. 

Subsequent firing tests on caliber .50 aircraft ma¬ 
chine gun barrels with 9-in. length, flanged-type, 
investment-cast liners of Stellite No. 21 inserted on 
a shrink-fit demonstrated outstanding performance 
of a different order of magnitude from that of stand¬ 
ard steel barrels, as is recounted in Chapter 22. 
Similar results were achieved with other rapid-fire 
small arms barrels with these liners. Under these con¬ 
ditions, Stellite No. 21 meets the criteria for a 
satisfactory bore-surface material. 

193 METALLURGY AND PROPERTIES 
OF STELLITE 

19,31 Chemical Composition of Stellites 

The stellites are cobalt-base alloys with 25 to 30% 
chromium, with a moderate amount (usually about 
6%) of molybdenum or tungsten, with small or mod¬ 
erate amounts of one or both of the metals iron and 
nickel, and with carbon contents varying from 0.2 to 
slightly over 1%. They are commonly cast alloys 
with manganese and silicon contents of 1 per cent or 
less. The nominal compositions of some stellites are 
given in Table 1. 

As the nickel or iron content is increased the ma- 
chinability of the alloy decreases. The stellites that 
contain molybdenum instead of tungsten are, in gen¬ 
eral, more ductile and softer than the tungsten-bear¬ 
ing alloys of the same carbon content. When the 
carbon content of a stellite is high, the alloy is strong¬ 
er, harder, and more brittle than when the carbon 
content is low. The effects of the chemical composi- 

e See Section 22.2.2 for details of this test. 


CONFIDENTIAL 





METALLURGY AND PROPERTIES OF STELLITE 


393 


Table 1 . Nominal percentage composition of stellites.* 



No. 21 f 

No. 

No. 23 

No. X-40 

No. 422-19 

No. 27 

Chromium 

28 

28 

25 

25 

25 

28 

Nickel 

2 

2 

2 

10 

16 

32 

Tungsten 

Molybdenum 

6 

5 

5 

7 

6 

5 

Cobalt 

Balance 

Balance 

Balance 

Balance 

Balance 

Balance 

Carbon 

(ca. 64) 

(ca. 65) 

(ca. 68) 

(ca. 58) 

(ca. 53) 

(ca. 35) 

0.25 

0.50 

0.40 

0.50 

0.40 

0.40 


* All of the cast stellites contain iron (maximum 2.5%), manganese (max 1%), and silicon (max 1%). 

t The tentative Army specifications^ for the composition of investment castings of cobalt-chromium alloy were about the same as this composition, 
t Stellite No. 22 is rolled Stellite No. 6. Stellite No. 6-2A is similar to Stellite No. 6 only with a lower carbon content. 


tion on liner performance are discussed later in this 
chapter (19.4.5 and 19.6). 

19 3 2 Crystallography and 

Structure of Stellite No. 21 

Investment-cast Stellite No. 21 has as its primary 
phase and matrix, a cobalt-rich, solid solution with a 
face-centered cubic lattice. The general structure is 
revealed in the photomicrograph shown as Figure 1. 
Numerous small, well-distributed areas of binary and 
ternary eutectic material are present. The binary 
eutectic is the M & C, solid solution type, where M may 
be cobalt, chromium, or molybdenum, and the ternary 
eutectic includes the phases Mq C, solid solution, and 
Cr 7 C 3 with part of this last carbide transformed to 
Cr 4 C. Small areas of a “pearlitic” constituent (from an 
analogy to the appearance of pearlite in steels) are 
present. This is believed to be a eutectoid formed by 
transformation of the cubic, cobalt-rich, solid solution 
to a hexagonal form with a simultaneous precipitation 
of Cr 4 C. 

The primary, cobalt-rich, solid solution phase of 
Stellite No. 21 is cubic as cast but can be largely 
converted into hexagonal metal (with some exsolu¬ 
tion of carbides) by heating at 800 C for about 50 hr. 
The cubic face-centered lattice of the cobalt-rich, 
solid solution is similar in unit cell size to the cubic 
modification of pure cobalt which is stable between 
about 400 C and 1000 C. The effect of the alloying 
elements chromium and molybdenum apparently 
raises the lower transition temperature and thus ex¬ 
tends the stability range of the low-temperature, 
hexagonal form to higher temperatures. 

Metallographic Examination 88 - 110 

After polishing, samples of Stellite No. 21 may be 
etched in a number of ways for metallographic ex¬ 


amination. Two methods have been found particu¬ 
larly useful. An electrolytic etch with dilute aqua 
regia reveals the general structure of the alloy while 
an alkaline potassium permanganate solution selec¬ 
tively stains the carbide constituents of the alloy. 

The structure varies only slightly from casting to 
casting and in different parts of the same casting. The 



Figure 1 . As-cast structure of a specimen from a 
typical investment-cast Stellite No. 21 liner: Electro¬ 
lytic aqua regia etch; 100X. (This figure has been taken 
from NDRC Report A-416.) 

rate of cooling during the casting process controls the 
grain size and the amount of the pearlitic constituent. 

A large number of liners of Stellite No. 21 were 
examined after firing tests. A photomicrograph of a 
fired liner is shown as Figure 7 in Chapter 16. Little 
or no change took place as a result of firing except at 
the immediate bore surface where a large number of 
surface cracks developed and there was evidence of 


CONFIDENTIAL 













394 


STELLITES AND OTHER COBALT ALLOYS 


age-hardening. Even after a severe firing test the 
surface cracks, although very numerous, were quite 
shallow and appeared at random and only occasion¬ 
ally at grain boundaries. 

X-Ray Examination 88 

X-ray examinations of Stellite No. 21 were made 
to see whether there were any peculiarities in the 
structure of investment-cast liners as compared with 
that obtained by other casting processes and also to 
observe any effects of heat-treatment or mechanical 
working on structure. Examination of liners prepared 
by investment-casting and by dry and green sand 
casting showed the investment-cast metal has much 
larger cubic crystals and that there is more preferen¬ 
tial orientation of crystals as a result of this casting 
process. 

Examination of metal at the bore surface of liners 
after severe firing tests showed cubic metal with a 
fine particle size. A few very faint lines were present 
on the x-ray photographs but could not be identified. 
Thus the essential change at the bore surface as a re¬ 
sult of firing was one of crystal size and not one of a 
chemical nature. The chemical inertness of this alloy 
towards hot powder gases makes it very erosion- 
resistant unless loss by melting occurs. 

By x-ray studies of samples heated to various tem¬ 
peratures and for various periods of time, the inver¬ 
sion temperature of hexagonal, cobalt-rich solution 
to cubic, cobalt-rich solution in this alloy was found 
to be approximately 800 C. This inversion is very 
sluggish; while pulverized metal heated at 800 C 
transformed in about 2 hours, solid pieces had not 
entirely transformed in 15 days. Any mechanical 
working at room temperature, such as cutting, filing, 
or machining, caused the surface layer to invert to 
the hexagonal form. 

While metallographic techniques do not readily 
distinguish the hexagonal metal from the cubic, x-ray 
examination quickly identifies the hexagonal phase. 

19,3,3 Effects of Heat Treatment 
on Structure 

Aside from age-hardening, the most noticeable ef¬ 
fect of heat treatment on the structure of Stellite No. 
21 is a change in the amount of the “pearlite.” 88 
When specimens of investment-cast metal are heated 
to 1300 C and held there for 15 minutes the pearlite 
dissolves. It forms again on cooling from 1300 C to a 


temperature between about 1050 and 950 C, and the 
amount formed depends on the cooling rate and con¬ 
sequently on the time the specimen is held in this 
range. 

19 3,4 Effects of Carbon Content on Structure 

The effects of variations of carbon content in Stel¬ 
lite No. 21 in the range from 0.08 to 0.58% carbon on 
the as-cast structure and hardness were studied. In 
this range both high and low carbon contents sur- 
press the formation of pearlite. Maximum pearlite 
developed in the alloy with 0.35% carbon. 

A similar series with normal carbon content (about 
0.25%) but with iron contents ranging from normal 
(2.5% or less) to as high as 20% were studied. Iron 
had no effect on the structure in either the as-cast or 
the heat-treated condition. There was no softening 
effect on hardness at room temperature, but such 
additions did decrease the hot-hardness. 

19.3.5 Work- and Age-Hardening 

The effectiveness of Stellite No. 21 as a liner ma¬ 
terial results not only from its hot-hardness and re¬ 
sistance to chemical attack by powder gases, but also 
from its susceptibility to work- and age-hardening 
and its retention of such added hardness during and 
after exposure to high temperatures. This conclusion 
is based on an extensive series of measurements. 88 An 
intense cold-working of the surface with resultant 
work-hardening occurs even during the machining 
operations (boring and rifling) on liners, and addi¬ 
tional work- and age-hardness is imposed by firing 
stresses. Cold-worked Stellite No. 21 age-hardens 
much more rapidly than the unworked metal and also 
reaches a higher ultimate hardness. 

19.3.6 Volume and Dimensional Stability 

Liners of Stellite No. 21 in machine gun barrels 
have been found to undergo small dimensional 
changes as a result of firing. With flanged-type liners 
inserted with a shrink-fit into a recessed caliber .50 
aircraft machine gun barrel, the length decreased 
during firing leaving a small gap at the forward joint, 
and there was a small decrease in bore diameter with 
an accompanying increase in muzzle velocity after 
the first burst of fire. These dimensional changes were 
probably caused for the most part by stresses during 
firing, with some slight changes caused by structural 


CONFIDENTIAL 



METALLURGY AND PROPERTIES OF STELLITE 


395 


changes in the liner alloy. Such dimensional changes 
have no deleterious effect on the performance of 
machine gun barrels. 

Past experience with stellite alloys had shown that 
small dimensional changes resulted from upsetting or 
plastic deformation when a piece of metal was pre¬ 
vented from expanding freely during heating because 
of a heavy shrink-fit. Since phase changes in the alloy 
were known to occur, a series of very careful experi¬ 
ments 88 were made to evaluate volume stability. No 
measureable changes were found after heating at 700 
C. Between 800 C and 1000 C a permanent shrinkage 
of 0.00086 inches per inch was observed. This change 
was slow (a matter of an hour or hours depending on 
the temperature). At 1200 to 1250 C dimensional 
changes were small and depended much on the pre¬ 
vious history of the sample. The changes varied from 
a small permanent expansion to a small permanent 
shrinkage. Chilling in liquid air caused no permanent 
dimensional changes. None of these changes were of 
great enough magnitude to account for the decrease 
in liner length during firing; such decrease in liner 
length must have resulted from bore stresses. 


19 3 7 Working Properties for 

Forging and Rolling 

Stellite No. 21 can be hot-forged or hot-rolled. 
Because of its high-temperature strength, hot-work¬ 
ing is much more difficult than with steels or other 
conventional materials. The essential requirements 
for the hot-rolling of Stellite No. 21 are the mainte¬ 
nance of a moderately high temperature (about 1190 
C) and a comparatively small reduction at each pass. 
During rolling, numerous reheatings and rolls that 
can apply high pressures are required, but sheet or 
strip of limited sizes can be produced in almost any 
gauge desired. The most desirable temperature range 
for forging is about 940 to 1070 C. Metal worked by 
forging or rolling shows the expected amount of grain 
refinement. Such metal is very hard and usually re¬ 
quires heat-treating (normalizing) at about 1150 C. 

Stellite No. 21 cannot be economically pierced and 
made into seamless tubing. For this reason castings, 
described in Section 19.4, rather than forgings, were 
used for liners. To obtain tubes of well-worked metal 
for tests, however, solid forged bars were drilled to 
the proper bore diameter. It is likely that welded 
tubing can be made from rolled strip and experiments 
on this process were in progress when the experi¬ 


mental work of NDRC was terminated. This process 
should be investigated further. 

19 3,8 Physical and Thermal Properties 

The available data on Stellite No. 21 may be sum¬ 
marized as follows: 

Hardness and Hot-Hardness. Various samples of 
Stellite No. 21 from liners showed hardness values 
between 331 and 387 VPN with 30-kg load at room 
temperature. Most liners tested about 26 on the 
Rockwell C scale. Measurements of the hot-hardness 
of investment-cast Stellite No. 21, with a 10-kg load, 
and corresponding measurements on WD 4150 ma¬ 
chine gun steel are shown in Table 2. This ability to 


Table 2. Hardness of Stellite No. 21 at different tem¬ 
peratures compared with that of WD 4150 steel. 


Temperature 

(C) 

Stellite No. 21 
hardness 
(VPN) 

WD 4150 steel 
hardness 
(VPN) 

20 

314 

290 

500 

275 

241 

600 

242 

140 

700 

206 

50 

800 

176 

36 

20 (after test) 

327 

202 


retain a large part of the room-temperature hard¬ 
ness at elevated temperatures is an outstanding prop¬ 
erty of the stellites. 

Specific Gravity and Density. The specific gravity at 
20 C is 8.30. Values for density at intervals between 
20 C and 800 C are given below under thermal co¬ 
efficients. 

Melting Range. This alloy does not have a sharp 
melting point but has a fusion range over which 
liquid and crystals coexist. This fusion range has not 
been located accurately and is influenced by varia¬ 
tions in chemical composition within specification 
limits for this alloy. The melting and softening ranges 
probably lie somewhere between about 1280 C and 
1350 C. Complete melting may even occur somewhat 
higher than 1350 C and values as high as 1396 C have 
been alleged. The melting interval needs further 
study. 

Thermal Coefficients. Temperature distribution in 
caliber .50 machine gun barrels with and without a 
stellite liner has already been discussed briefly in Sec¬ 
tion 5.4.4. The thermal coefficients of Stellite No. 21 
as compared with machine gun steel are given here in 
Table 3. Stellite is a poorer heat conductor than steel. 


CONFIDENTIAL 







396 


STELLITES AND OTHER COBALT ALLOYS 


Table 3. Thermal coefficients of Stellite No. 21 and of SAE 4150 steel for comparison (in parentheses). 106 


Temp 

(C) 

Density 

(gm/cm 3 ) 

Volume 

specific 

heat 

(cal/cm 3 /°C) 

Conductivity 
(cal/cm sec/°C) 

Diffusivity 

(cm 2 /sec) 

Heat 

conduction 
constant 
(cal/cm 2 sec*/°C) 

20 

8.30 

0.96 




200 

8.24 <7.79) 

1.00 (0.98) 

0.035 (0.100) 

0.035 (0.102) 

0.19 (0.31) 

300 

8.20 

1.03 

0.039 

0.038 

0.20 

400 

8.17 

1.06 

0.043 

0.41 

0.21 

500 

8.12 

1.09 

0.047 

0.043 

0.23 

600 

8.09 (7.65) 

1.13 (1.35) 

0.051 (0.081) 

0.045 (0.060) 

0.24 (0.33) 

800 

8.01 

1.23 





In the case of a single round in a stellite-lined barrel, 
this results in a lower heat input per round with 
higher bore-surface temperatures but lower tempera¬ 
tures at depth as compared with a regular steel bar¬ 
rel. A further outcome of this difference in thermal 
properties is that in a continuous burst the deep 
swaging temperatures f in the region of the liner are 
not very different than those for a regular steel barrel 
for 200 rounds. For longer bursts the deep swaging 
temperatures rise more rapidly in steel than in stel¬ 
lite. Thermal effects in gun barrels during firing are 
discussed in detail elsewhere. 106 

Dilatometer curves made with an optical, differ¬ 
ential dilatometer between room temperature and 
1000 C showed no thermal arrests and the heating 
and cooling curves were nearly identical. 80 

19 3 9 Mechanical Properties 

Considerable data on the tensile properties, hard¬ 
ness, impact strength, modulus, and ductility of Stel¬ 
lite No. 21 both at room temperature and at elevated 
temperature were obtained during the stellite-liner 
development. The effects of method of casting, heat- 
treatment and mechanical working on these proper¬ 
ties were evaluated. 80 Stellite No. 21 has excellent 
strength and hardness, as well as reasonably high 
ductility and impact resistance, both at room tem¬ 
perature and at the elevated temperatures encount¬ 
ered in machine gun barrels during firing. Informa¬ 
tion on the mechanical properties of the investment- 
cast metal in the as-cast condition was obtained by 
mechanical tests. The data for tests at two tempera¬ 
tures are given in Table 4. Data on SAE 4150 steel 
are given for comparison. The test bars were 1-in. 
gauge by 0.24-in. Hot-hardness values were given 
in Table 2. 


f Temperature at in- from bore surface. 106 


Measurements 88 of modulus of elasticity (Young’s 
modulus) showed decided directional properties in in¬ 
vestment-cast metal. Because of the very large grain 
size the entire cross section of a 0.25-in. tensile speci¬ 
men may consist of a single grain. Moduli ranging 
from 23 million to 40 million psi were measured. It is 
interesting to note that such nonuniformity of mod¬ 
ulus had no noticeable effect upon life and perform¬ 
ance in firing tests. It should be remembered that all 
cast (and heat-treated) metals, such as steels, brasses, 
etc., have a rather indefinite elastic modulus, spread 
over a wide range of experimental values. Only 
wrought metals have fixed values for their moduli. 

19310 Machinability 

Stellite No. 21 can be machined readily with prop¬ 
er tooling. In spite of much apprehension on the 
part of machine gun barrel manufacturers, no serious 
problems were encountered in the machining of stel¬ 
lite liners in production. By the use of the proper car¬ 
bide-tipped tools and cuts and feeds 119 determined in 
the development work, 80 it was found that it was 
much easier to hold Stellite No. 21 to close tolerances 
on dimensions than machine gun steel (WD 4150). 

19311 Availability 

Stellite No. 21 contains more than GO per cent of 
cobalt and about 28 per cent of chromium. During 
the war the supply of both metals was critical. Fer- 
rochrome, which was less critical than pure chromi¬ 
um, was used to supply a part of the chromium. 

Only because there was an adequate stock pile of 
cobalt, were all of the stellite liners needed for machine 
gun barrels obtained. Practically no cobalt ore is pro¬ 
duced in the United States. The world supply comes 
largely from the Belgian Congo. During times of war, 
shipping conditions may be such that this supply will 


CONFIDENTIAL 














UTILIZATION OF STELLITE NO. 21 397 


Table 4. Results of mechanical tests on investment-cast Stellite No. 21 in as-cast condition* and on SAE 4150 steel.f 

Sample 

Temp 

Tensile 

strength 

(psi) 

Yield 

point 

(psi) 

Breaking 

strength 

(psi) 

Elon¬ 

gation 

(%) 

Reduction 
of area 

(%) 

Hard¬ 

ness 

(Rc) 

Stellite 

Room 

90,000 

65,000 

123,000 

23.5 

23 

29 

Steel 

Room 

135,000 

116,000 


21 

57 


Stellite 

1500 F 

71,000 

45,000 

84,000 

30 

24 


Steel 

1200 F 

45,000 

20,000 

— 


38 

92 



* Tests made at Crane Co. and results given in Report R on Contract OEMsr-629. The Charpy impact-resistance of this material varies from 9 to 


13 ft-lb at 80 F and from 11 to 13 ft-lb at 1000 F. 

t Tests made at Crane Co. Results given in letter to Chief, Division 1, August 7, 1944. 


not be available. Cobalt occurs as an accessory in the 
Canadian nickel ores but is usually not extracted be¬ 
cause of cost considerations. If stellite is to be used in 
large amounts for gun liners, attention must be given 
to an assured, large cobalt supply in wartime or a 
large stockpile must be available. 

194 UTILIZATION OF STELLITE NO. 21 
AS A LINER OR LINING MATERIAL 

19,41 Introduction 

In order to insure the most effective utilization of 
the remarkable erosion-resistance and good mechan¬ 
ical properties of Stellite No. 21, it was essential to 
evaluate the various possible methods of fabricating 
this alloy and of applying it to gun bores of various 
sizes with respect to the feasibility of the methods 
and to the effects of various factors on the perform¬ 
ance of machine gun barrels containing stellite liners. 

19 4 2 Casting Methods 

The first liners of Stellite No. 21 tested were pre¬ 
pared in investment molds. These investment- or pre¬ 
cision-castings have proved very satisfactory and 
nearly one-half of a million liner castings were made 
for insertion in machine gun barrels of the type de¬ 
scribed in Chapter 22. This casting method was ex¬ 
pensive as compared with certain other ones, but the 
overall cost was low, because these precision castings 
required a minimum of grinding and machining oper¬ 
ation to complete a liner for insertion. 

Firing tests 80 were performed on caliber .50 machine 
gun liners cast by the following methods: (1) green¬ 
sand-mold casting, (2) dry-sand-mold casting, (3) 
graphite-mold casting, and (4) centrifugal casting in 
refractory molds. When proper melting and deoxida¬ 
tion techniques were used and the rate of cooling of 
the castings was controlled to approach the slow rate 


characteristic of casting in hot investment-molds, 
liners prepared by dry sand casting and by centri¬ 
fugal casting in refractory molds were found to be 
equal in gun performance to those made by invest¬ 
ment-casting. However, both of the former types of 
casting, especially the centrifugal one, require con¬ 
siderably more grinding and machining to finish the 
liners for insertion. In green sand molds it was diffi¬ 
cult to get sound castings, and graphite-mold castings 
showed a much greater “wear” during firing tests than 
investment-castings. The centrifugal thermit-casting 
method 90 and other centrifugal melting techniques 82 
were being studied but no liners made by these meth¬ 
ods were tested before termination of the experi¬ 
mental program. 

One important requirement for any stellite-liner 
casting is freedom from unsoundness. To insure qual¬ 
ity and performance, it was found necessary to radio¬ 
graph all liner castings in two directions at right 
angles. Although the soundness of castings could be 
determined more definitely when radiographed after 
rough turning and boring, it was found necessary in 
the interests of speed and economy of production to 
radiograph before machining. A series of firing tests 
was conducted to correlate the appearance of the 
radiographs with firing performance. It has been de¬ 
finitely established that radiographically sound cast¬ 
ings insure satisfactory firing performance. In addi¬ 
tion, the effects of variations of core materials and 
changes in the technique of investment casting were 
correlated with performance on firing. All of these 
studies insured the reliability of liner performance in 
Service weapons. In conclusion, it may be stated that 
investment casting was adopted as the standard 
method of preparing Stellite No. 21 because of a pre¬ 
viously well-established practice. It has been suffi¬ 
ciently well demonstrated, however, that liners pre¬ 
pared from any radiographically sound casting per¬ 
form satisfactorily; thus, the selection of a casting 
method depends simply on economic considerations. 


CONFIDENTIAL 






398 


STELLITES AND OTHER COBALT ALLOYS 


19 4 3 Infusion, Incasting, and 

Torch Deposition 

Advantages of Fusion-Bonded Linings 

It was expected that if Stellite No. 21 could be at¬ 
tached to steel with a satisfactory fusion- or casting- 
bond between the tw T o metals, a bore lining of stellite 
prepared in this way would present the following ad¬ 
vantages : 

1. It w ould permit a saving of critical and expensive 
metal by elimination of the metal of the flange and 
would also permit the application of thinner layers 
than is practical with castings. (It is difficult to main¬ 
tain tolerance during machining and rifling opera¬ 
tions in thin-walled castings owdng to distortion.) 

2. It would eliminate the necessity for a retaining 
nut behind the breech end of the liner, thus saving 
machining operations. 

3. It would make full-length linings practical. (The 
preparation and mechanical insertion of full-length 
liners is very difficult as pointed out in Section 
19.4.6.) 

4. It would be a practical method of applying a 
stellite bore surface to the w r hole or part of the bore of 
large gun tubes. (It is difficult to make large castings 
that are sound and free from blow holes, shrinkage 
'cavities, or other imperfections, and the final prepa¬ 
ration of the liner casting and recess in the gun tube 
to receive it involves expensive and time-consuming 
machine operations.) 


Applications of Fusion-Bonded Linings 

Stellite No. 21 was applied as full-length or partial- 
length linings in caliber .50 aircraft machine gun bar¬ 
rels by several methods of infusion and incasting. 



Figure 2. Longitudinal section of Stellite No. 21 rod 
fused in steel sleeve with steel plug at end of sleeve: 
Steel etched in nital; Stellite No. 21 etched electro- 
lytically in sodium cyanide solution. (This figure has 
been taken from a report on Contract OEMsr-629.) 


These same methods were used later to prepare re¬ 
placeable steel liners with a stellite lining for tests in 
larger gun tubes, as described in Section 33.1.2. Infu¬ 
sion consists of the fusion of a casting in place in a re¬ 
cess in a steel tube. Incasting consists of filling a 
steel tube (w T ith or without a core) with molten stel¬ 
lite so that the steel and stellite are bonded. A bond 
produced.by infusion is shown in Figure 2. 

Composite caliber .50 liners for test purposes were 
prepared by the several methods listed below and 
were subjected to firing tests to evaluate the behavior 
of the bond and the lining material. When suitably 
prepared, the bond in all cases was found to be excel- 



Figure 3. Caliber .50 barrel, stellite-lined by progres¬ 
sive static infusion and fired 2,008 rounds according 
to the CGL-350 schedule. (This figure has appeared as 
Figure 15 of NDRC Report A-417.) 

lent according to metallographic evidence and the 
linings survived all the rigors of firing. Figure 3 shows 
a fired caliber .50 barrel which had a stellite lining 
infused by the first method given below. 

Progressive Static Infusion™ A series of short cast¬ 
ings was progressively melted in place in a vertical 
steel tube with a ceramic or other core by use of a 
moving induction coil. A diagram of the furnace 
assembly is shown as Figure 4. The apparatus for 
preparing a lining for a 37-mm gun (Section 33.1.2) 
by this method is shown in Figure 5. 

Centrifugal Infusion or Incasting. Pieces of solid 
stellite were melted in a spinning steel tube, or molten 


CONFIDENTIAL 








UTILIZATION OF STELLITE NO. 21 


399 


stellite was poured into a steel tube which was spun 
until the stellite had solidified. 125 In addition, stud¬ 
ies 90 were made of centrifugal, thermit incasting, but 
no liners for test were prepared by this method before 
termination of the experimental program. 

Cast-Welding 88 Molten stellite was poured into the 
space between a heated steel tube and a ceramic or 
other core. 

Vacuum Incasting . 90 Molten stellite was sucked up 
or down into the evacuated space between a steel tube 
and a ceramic or other core. A diagram of the mold 
arranged for down-sucking is shown as Figure 6. 

Torch Deposition . 80 A stellite surface was applied as 
a bore surface on two halves of a flanged, steel cal¬ 
iber .50 liner by torch deposition using an oxyacety¬ 
lene welding torch. The behavior of the metal and 
bond during firing was excellent. Torch deposition is 
not practical for small bores such as caliber .50 but 
shows much promise for lining large gun tubes. 

Three processes of torch deposition were studied; 
in two, welding was done with oxy acetylene and 
atomic-hydrogen torches, and the third employed a 
metallic arc. The atomic-hydrogen-torch welding 
process appears to be the most practical method of 
lining large-caliber gun tubes because of the possibil¬ 
ity of automatic or semi-automatic application. The 
oxyacetylene-torch method must be considered un¬ 
suitable until a flux is developed which will permit a 
neutral flame adjustment to yield deposits free from 
pickup of excess carbon. The metallic-arc welding 
process appears promising, especially as a means of 
repairing surface defects in stellite deposits. Studies 
of these processes should be continued. 

Necessary Precautions in Fusion-Bonding 

In all of the infusion or incasting methods care was 
necessary to prevent serious contamination of the 
stellite by steel, especially in those methods which in¬ 
volved pouring molten metal along a steel surface or 
centrifugal spinning of molten stellite against steel. 
The particular problems involved in preparing full- 
length linings are discussed in Section 19.4.6, and the 
effects on performance of contamination by steel are 
discussed in Section 19.4.5. 

19 4 4 Other Methods of Applying Stellite 

The application of Stellite No. 21 to gun bores as 
cast liners has already been described in Section 
19.4.2, liners bored from forged rods in Section 19.3.7, 


composite liners or linings prepared by infusion, in¬ 
casting, or torch deposition in Section 19.4.3, and the 
possibilities of its application as seamless tubing or 
welded tubing made from rolled strip has been dis¬ 
cussed in Section 19.3.7. In addition to these, three 



Figure 4. Diagram of furnace assembly ready for 
progressive static infusion of Stellite No. 21. (This figure 
has appeared as Figure 4 of NDRC Report A-417.) 


other methods of applying stellite linings or coatings 
to gun bores were explored as follows: 

Sprayed Coatings . 90 Stellite No. 21 was applied to 
steel surfaces by the use of a “spray-gun.” The pro¬ 
duction of an adherent deposit requires a rough steel 
surface. In the case of steel tubes with an internal 
diameter of less than 13 ^ in., spray methods were im¬ 
practical. Even the best spray-coats of stellite showed 
lamination and porosity, and tensile test specimens 
cut from thick spray coats showed zero elongation 
and a strength of less than 9,000 psi. The mechanical 


CONFIDENTIAL 

















































































400 


STELLITES AND OTHER COBALT ALLOYS 



Figure 5. Induction-heating furnace for infusing 37-mm blanks. (This figure has appeared as Figure 2 of NDRC 
Report A-417.) 


CONFIDENTIAL 








































































UTILIZATION OF STELLITE NO. 21 


401 


properties of such deposits make them unsuitable as 
bore-surface materials. Spraying might be used as a 
method of applying a layer of stellite prior to an in¬ 
fusion or torch-welding process. 

Coatings by Thermal Explosion . 88 Metal can be 
vaporized by heat and such vapor can be deposited 
to form a coating. Stellite wires were exploded in a 
vacuum by an electrical discharge and the metal was 
deposited on the walls of a steel tube. The rate of de¬ 
position was extremely slow and the coating was un¬ 
satisfactory unless its adherence was improved by a 
diffusion treatment. The process is expensive and 
thoroughly impractical. 

Electroplating and Diffusion . 88 It is possible to apply 
the component metals of Stellite No. 21 in the form 
of a series of thin electrodeposits and then convert 
these layers into the alloy by a high-temperature dif¬ 
fusion process. It was found that the plates could be 
applied and subsequently alloyed. These experiments 
were not carried along far enough to estimate the 
commercial feasibility of such a method. 

19 4 5 Effects of 

Various Factors on Performance 

An extensive series of firing tests 80 was conducted 
on stellite-lined caliber .50 aircraft machine gun bar¬ 
rels in order to evaluate the importance of the various 
factors which control the performance of this barrel 
with a breech liner of Stellite No. 21. In addition to 
the evaluation of the effects of casting methods and 
infusion, incasting, and torch deposition methods al¬ 
ready described in Sections 19.4.2 and 19.4.3, the 
effects of the following additional factors were studied. 

Hardness 

The effects of variations in hardness in regular, 
9-in. investment-cast liners was studied. Two thou¬ 
sand liners were checked for hardness and with few 
exceptions were found to range in hardness from 20 to 
30 Rockwell C with an average of 25.8. Two liners 
with the lowest hardness and two with the highest 
were selected for test, and performance was compar¬ 
able to that of liners of normal hardness. 

A study of liners with work-hardened bores was in 
progress, but the experiments were not completed be¬ 
fore termination of the program. 

Liners with hardened bores produced by the draw¬ 
rifling processes 435 were subjected to firing tests. The 
results were inconclusive because the rifling was not 


of standard dimensions and frequent stoppages oc¬ 
curred during firing. 

Nitriding 


During the research program to improve further 
the barrel containing a stellite liner by chromium¬ 
plating the steel bore ahead of the liner, which is the 



Figure 6 . Mold for vacuum incasting of Stellite No. 
21. (This figure has appeared as Figure 2 of NDRC 
Report A-418.) 


CONFIDENTIAL 























































402 


STELLITES AND OTHER COBALT ALLOYS 


subject of Chapter 24, nitriding of the steel bore prior 
to plating was tried. To simplify the procedure the 
nitriding was performed with the stellite liner in 
place. Nitriding had little, if any, effect on Stellite No. 
21 and when barrels containing stellite liners were 
nitrided and fired, no difference in performance 81 was 
noted as compared with regular stellite-lined barrels. 

Decarburization 

Early in the program it was felt that the presence 
of eutectoid at the grain boundaries of stellite might 
be a source of weakness. This material might suffer 
preferential melting or attack and cause cracking of 
the bore surface. Several liners were decarburized 88 
at 1175 C for 5 hr in dry hydrogen. All of these liners 
gave inferior performance except one which had been 
decarburized before machining. This liner gave per¬ 
formance comparable to regular stellite-lined barrels. 

Hot-Working 

Since much higher strength and hardness and 
equivalent ductility and better impact resistance 
were obtained in laboratory tests on metal mechan¬ 
ically worked by rolling or forging, as compared with 
cast metal, two caliber .50 liners were prepared by 
boring cast rods which had been forged to the extent 
that the diameter was decreased from 2 in. to 1 in. 
Firing tests on these liners showed performance equal 
to but not superior to that obtained with regular 
investment-cast liners. 

Chemical Composition 

Firing tests on stellites other than No. 21, which 
are described later (Section 19.6), indicated that the 
chemical composition of the liner alloy was an im¬ 
portant factor in determining performance. To insure 
reliability of performance, exploratory studies of the 
effects of varying carbon contents, varying iron con¬ 
tents, and varying molybdenum, iron, and carbon 
contents of Stellite No. 21 were made. 

A series of investment-cast liners was prepared 
with carbon contents varying by steps from 0.08 to 
0.58%. Two liners of each composition were subjected 
to firing tests. The normal carbon content of Stellite 
No. 21 is 0.20 to 0.35%. The firing tests showed that 
liners with the normal range of carbon content per¬ 
formed in a very satisfactory manner and that the 
percentage could be increased to as much as 0.58 


without any impairment in performance. When, how¬ 
ever, the carbon was as low as 0.08%, performance 
was distinctly inferior. 

The effect of replacing some of the cobalt with iron 
and thus increasing the iron content beyond the 
specified maximum of 2.5% was of interest because 
of a critical shortage of pure chromium, which might 
require the use of ferrochromium in the production of 
liner castings, and because increased iron content 
through pickup from steel is often obtained during 
the infusion, incasting, and torch deposition methods 
already described in Section 19.4.3. Firing tests on 
investment-cast liners prepared with and without 
ferrochromium and with iron contents varying by 
steps from 3.3 to 20.15% showed that the maximum 
iron content that can be tolerated without seriously 
affecting performance is about 7%. 

A few exploratory experiments were made in which 
two of the constituents were increased and the cobalt 
decreased. Firing tests on investment-cast liners with 
the percentages of the constituents varied as follows: 

8.60% Fe and 8.90% Mo 
9.85% Fe and 0.47% C 
7.10% Fe and 0.46% C 

showed, respectively, equal, decidedly inferior, and 
slightly inferior performance as compared with regu¬ 
lar liners of Stellite No. 21. 

Heat-Treatment 

A systematic study was made of the effect of heat 
treatments on the physical properties, microstruc¬ 
tures, and firing performance of both investment-cast 
and sand-cast Stellite No. 21. All the tensile proper¬ 
ties, hot-har^ness, and microstructure can be varied 
by heat-treatment within a rather wide range. A 
given sample can be made ductile with a relatively 
low hardness, while another can be made very hard 
with relatively low ductility and impact resistance. 
There is no significant difference in physical proper¬ 
ties of the metal whether prepared by investment- or 
sand-casting methods. 

A series of investment-cast liners was subjected to 
selected heat treatments based on the above observa¬ 
tions and subjected to firing tests. None of the heat 
treatments improved the performance over that of 
the as-cast material. Some heat treatments had no 
effect while others definitely reduced the performance. 

One important observation was that when an in¬ 
vestment-casting was subjected to the heat-treat- 


CONF1DENTIAL 




UTILIZATION OF STELLITE NO. 21 


403 


ment normally given to gun steel (oil quenching from 
1550 F and drawing at 1000 F), there was no effect 
on its performance. Thus, composite Stellite No. 21- 
steel assemblies made by infusion or incasting (see 
Section 19.4.3) can be heat-treated to develop the 
desired properties in the steel without impairing in 
any way the performance of the stellite. 

Firing tests of investment-cast liners of Stellite No. 
21 that had been heat-treated to develop a maximum 
and a minimum amount of the pearlitic structure 
showed that neither of these heat treatments was 
advantageous. 

Method of Insertion 

In the preparation of cast-stellite liners for inser¬ 
tion in steel machine gun barrels on a production 
basis, it is desirable to allow as large tolerances as pos¬ 
sible on the amount of interference between the liner 
and barrel recess for a shrink-fit. Too small an inter¬ 
ference may allow the liner to rotate during firing, 
while too large an interference may cause both exces¬ 
sive bore-constriction during firing and difficulties in 
assembly. Firing tests established a maximum dia¬ 
metrical interference of 0.001 in. on the portion of the 
liner with small diameter. This may be increased to 
0.002 in. on the shoulder or flange of the liner. 

There was some evidence that the immersion of 
liners in liquid nitrogen (either before or after machin¬ 
ing or in the process of insertion) slightly enhanced 
the performance and that liners inserted by heating 
the barrel only, without cooling the liners, gave 
somewhat inferior performance. This tentative con¬ 
clusion should be checked by more comparison-fir¬ 
ing data. 

Liner Length 

Nonplated Steel Ahead of Liner. Six-inch, nine-inch, 
and twelve-inch (overall length) liners were subjected 
to firing tests in caliber .50 aircraft machine gun bar¬ 
rels. Severe erosion of the steel bore ahead of a 6-in. 
liner occurred and seriously affected performance. 
The 12-in. liner was unsatisfactory because of a dan¬ 
gerous weakening of the barrel near the forward joint 
after a gap had been produced by a slight shrinkage 
in liner length during firing. The 9-in. liner gave very 
satisfactory performance. The problems of full length 
liners are discussed later in Section 19.4.6. 

Plated Steel Ahead of Liner. Chromium-plating of 
the steel bore ahead of a stellite liner considerably 


enhances the performance of the machine gun barrel. 
This combination of a 9-in. liner with plate and its 
Service applications are described later in Chapter 
24. Studies were made of the performance of barrels 
with 4-in., 6-in., and 9-in. liners of Stellite No. 21 
with tapered (choke-muzzle) chromium-plate ahead 
of the liner. Barrels with a 4-in. liner plus the plate 
were equivalent in performance to barrels with a 
9-in. liner without plate; those with 6-in. liners 
plus plate were somewhat superior to the latter; 
and those with 9-in. liners plus plate were very much 
superior. 

Attention should be called at this time to the fact 
that, in order to get full benefit from a short erosion- 
resistant liner of Stellite No. 21 with or without an 
erosion-resistant plate (chromium plate) ahead of the 
liner, the barrel weight should be increased, the con¬ 
tour (distribution of steel) should be changed, and 
the composition of the barrel steel should be changed. 
These matters are taken up fully in Chapter 24. 

Wall Thickness of Liners 

The J/g-in. wall thickness was selected for the liners 
on the basis of the strength of the liner and its resist¬ 
ance to deformation during handling and during the 
various grinding, machining, boring and rifling oper¬ 
ations necessary to prepare a liner casting for inser¬ 
tion. With considerable care it was possible to prepare 
two liners with a ^-in. wall thickness. Their firing 
performance was entirely comparable to regular 
liners with 3^8-in. walls. This result is of considerable 
interest in connection with the performance of stellite 
when applied to gun bores as thin linings by infusion, 
incasting, or torch deposition. 

Plated Liners 

If a bore surface of Stellite No. 21 with its excellent 
resistance to the swaging impact of projectiles could 
be protected from surface melting, it might be pos¬ 
sible to extend the useful range of application of stel¬ 
lite to hypervelocity conditions with the resultant 
high bore-surface temperatures. Attempts were made 
to protect stellite bore surfaces with deposits of the 
high-melting, erosion-resistant metals chromium, 
molybdenum, and tungsten. A number of investment- 
cast caliber .50 liners were rifled oversize to allow for 
such coatings, and attempts were made to apply 
satisfactory deposits of chromium by electroplating 
and of tungsten and molybdenum by pyrolytic plat- 


CONFIDENTIAL 



404 


STELLITES AND OTHER COBALT ALLOYS 


ing, as described in Chapter 21. Electroplated chro¬ 
mium was poorly adherent on stellite surfaces and no 
firing tests were made. Diffusion bonding to increase 
adherence of this plate should be studied. Pyrolytic 
coatings of tungsten and molybdenum on stellite 
liners were subjected to firing tests and were unable 
to withstand even a mild schedule without spalling 
and deformation. 

19,4,6 Preparation and Testing of 
Full-Length Liners or Linings 

Linings Prepared by Static Infusion 

Caliber .50 aircraft machine gun barrels with par¬ 
tial length and full-length linings of Stellite No. 21 
were prepared by the method of progressive static in¬ 
fusion. 89 This method provides barrel blanks with 
partial or full-length linings ready for finishing as one- 
piece barrels and eliminates liner insertion and the 
necessity of the retaining nut behind the liner at the 
breech end of the barrel. Particular problems were 
encountered in the preparation of satisfactory full- 
length linings. Troublesome coring problems were en¬ 
countered and solved. Preliminary cost analysis stud¬ 
ies indicated that it should be possible to manufac¬ 
ture such barrels at substantially lower costs than 
barrels with inserted 9-in. liners. 125 

When barrels were completed and subjected to fir¬ 
ing tests 80 it was found that, unless the stellite lining 
was of very uniform thickness and was perfectly con¬ 
centric with the outside diameter of the steel barrel 
wall, the shot-pattern moved off the target during fir¬ 
ing owing to the bimetallic effect and to reversible 
warping of the composite barrel as a result of heating 
during firing. Even when the lining was of uniform 
thickness and concentric, and dispersion was normal, 
performance was only equivalent to that of barrels 
with 9-in. long, investment-cast stellite liners inserted 
on a shrink-fit, and was inferior to that obtained with 
barrels containing a 9-in. liner plus tapered (choke- 
muzzle) chromium plate ahead of the liner. 

Liners in Series 

A 13-lb caliber .50 aircraft machine gun barrel 
was recessed to receive a nearly full-length bore- 
surface of Stellite No. 21 by inserting three invest¬ 
ment-cast liners. 80 The stellite extended to within 1- 
in. of the muzzle; the forward two liners had no 
flanged shoulders. During the firing test, one of the 


forward liners rotated and a 0.25-in. separation oc¬ 
curred between the forward two liners as a result of 
shrinkages in length during firing. There was no ten¬ 
dency for the shot-pattern to move off the target as 
with some of the full-length, infused linings previously 
described. Overall performance was unsatisfactory. 

“In-melted” Liners 

Towards the close of the experimental program it 
was possible to produce full-length liners of Stellite 
No. 21 which could be shrunk into standard or special 
caliber .50 aircraft machine gun barrels. They were 
prepared by a process of remelting a stacked series of 
short, tubular fillers of investment-cast stellite within 
an inductively heated, refractory-lined, graphite 
tube. 89 One of these liners before machining is shown 
here as Figure 7. Stress-relieving treatments, machin¬ 
ing procedures which would preserve the straightness 
and concentricity of the liners, and a proper insertion 
procedure were developed. Two barrels with these 
liners were ready for firing tests, which were not 
carried out, however, because of the termination of 
the experimental program. 

19 5 APPLICATION OF LINERS OR LININGS 
OF STELLITE NO. 21 TO GUNS 

19,5,1 Improved Machine Gun Barrels 

Experience of aircraft combat during the war in¬ 
dicated that erosion was limiting the performance of 
the caliber .50 aircraft machine gun. The application 
of short, breech liners of Stellite No. 21 to machine 
gun barrels led to remarkable increases in the length 
of life. Barrels containing stellite liners, the subject 
of Chapter 22, showed outstanding performance un¬ 
der firing conditions so severe that unmodified steel 
barrels were unable to withstand them. 

The combination of the stellite liner with tapered 
(choke-muzzle) chromium plate ahead of the liner 
gave further outstanding improvement in perform¬ 
ance particularly when the weight of the barrel, con¬ 
tour (distribution of steel) and barrel steel compo¬ 
sition were adjusted to enable maximum utilization of 
the erosion resistance of the liner and plate. The 
types of barrels that were thus modified and im¬ 
proved are discussed in Chapter 24. 

When double-base powder was used in some of the 
acceptance tests of improved caliber .50 aircraft bar- 


CONFIDENTIAL 



FIRING RESULTS ON STELLITES OTHER THAN NO. 21 


405 



Figure 7. In-melted, full-length, Stellite No. 21 liner; cast surface except at joints of original fillers. (This figure has 
appeared as Figure 2 of NDRC Report A-417.) 


rels, mentioned in Section 22.3.3, conditions ap¬ 
proached those of hypervelocity. The failure of stel¬ 
lite under these conditions is discussed in the next 
section. 

19 5 2 Hypervelocity Guns 

Two liners of Stellite No. 21 and two of Stellite No. 
22 were tested in the caliber .50 erosion-testing gun 
(Section 11.2.1) under conditions of hypervelocity at 
a slow rate of fire. 76 77 For one pair of tests, double¬ 
base powder (20% nitroglycerin) was used and for 
the other pair single-base (IMR) powder. With 
double-base powder, the bore surface of the stellite 
liners melted so rapidly that it was impossible to 
establish full pressure. With single-base powder, with 
which a muzzle velocity of 3,571 fps was achieved, 
only a small amount of local melting occurred at the 
bore surface. The latter tests were discontinued after 
500 rounds, at the end of which the rifling of the 
liners was still in very satisfactory condition. Figure 
6 of Chapter 16 shows the contrast between the bore 
surfaces of liners fired with these two kinds of 
powder. 

Experience with stellite liners in the caliber .60 
machine gun barrel showed that this alloy is marginal 
with respect to its use in hypervelocity guns. This 
gun, for which liners of Stellite No. 21 are being used 
regularly, has a muzzle-velocity of slightly over 3,500 
fps with IMR powder. In this particular application, 
a stellite liner, unlike steel, lasts long enough to 
furnish a useful gun-barrel life. However, metallo- 
graphic studies have shown that ultimate failure is 
the result of melting along surface cracks. 

These observations showed that an erosion-resis¬ 
tant material of higher melting point than the stellites 
was required for general use in hypervelocity guns. 
The development of such materials is described in the 
other chapters of Part “V. 


19 5 3 37-mm Cannon 

A partial bore lining of Stellite No. 21 has been 
successfully applied to the 37-mm gun tube, M3. This 
development is described in Section 33.1.2. 

196 FIRING RESULTS ON STELLITES 
OTHER THAN NO. 21 

The early firing tests on Stellite No. 6 and the 
reasons which led to the selection of Stellite No. 21 
have already been described in Section 19.2. Firing 
tests 80 have been made with caliber .50 liners pre¬ 
pared from other stellite-type alloys, the composi¬ 
tions of which are given in Table 1, to determine (1) 
whether any of these might be superior to Stellite 
No. 21 as a gun liner, or (2) whether any of them 
containing less of the critical metal cobalt might be 
equally suitable. None of the other stellite-type 
alloys gave performance better than that of Stellite 
No. 21. 

The following tungsten-bearing stellites were test¬ 
ed: No. 6, No. 22, No. 6-2A, No. 23, and No. X40. 
All showed inferior performance to Stellite No. 21 
and greater land “wear” during firing, and all of the 
liners except the one of Stellite No. 23 exhibited deep 
cracking and brittle failure at some stage of the firing 
tests. 

Two molybdenum-bearing stellites in which nickel 
had been substituted for some of the cobalt were 
tested as liners. Stellite No. 422-19 showed greater 
land wear. One investment-cast liner of Stellite No. 
27 showed performance equal to that of Stellite No. 
21, but a duplicate liner from the same casting-heat 
had no rifling left and the barrel showed a large 
velocity-drop after one continuous burst of 300 
rounds in the caliber .50 aircraft machine gun. When 
two additional liners of No. 27 were fired later they 
both showed complete obliteration of the rifling and 


CONFIDENTIAL 







406 


STELLITES AND OTHER COBALT ALLOYS 


a large velocity drop after one long burst of fire. This 
alloy is very difficult to cast because of shrinkage in 
the mold and is gummy and difficult to machine. 

197 COBALT AND HIGH-COBALT ALLOYS 
AS LINERS 

In vent-plug tests (Section 16.3.1) pure cobalt was 
slightly less resistant to powder-gas erosion (melting 
and chemical attack) than gun steel under the severe 
conditions of the %-in. vent test but under the milder 
conditions of the J^-in. vent test it was more resist¬ 
ant. 50 This result was confirmed by a test of a short, 
rifled caliber .30 liner of pure, swaged cobalt (Section 
16.3.6). However, although it exhibited resistance to 
powder-gas erosion and good ductility, pure cobalt 
had insufficient hot-hardness to resist deformation of 
the rifling by swaging impact of the bullets. 59 

In an effort to harden cobalt and if possible to 
increase its resistance to powder-gas erosion, binary 
cobalt alloys containing varying amounts of tungsten, 
molybdenum, or chromium were prepared as ingots 
and swaged into rods. These alloys eroded less than 
stellites in vent-plug tests. 59 

The alloy containing 93% cobalt and 7% tungsten 
showed excellent performance when tested as a liner 
in the caliber .50 heavy machine gun barrel. 81 Since 
these alloys contained more critical metal cobalt, 
were much more difficult to prepare than the easily 
cast Stellite No. 21, and showed no marked advantage 
in performance over the latter, their further develop¬ 
ment as liners was not pursued. 

19 8 ALLOYS CONTAINING LESS COBALT 
THAN MOST STELLITES 

1981 Introduction 

During the course of the stellite-liner development, 
difficulties arose because of the shortage of critical 
materials. Ordnance Department liaison officers urged 
Division 1, NDRC, to develop substitutes for Stellite 
No. 21 with less cobalt or if possible without any, in 
order to save the very critical cobalt, part of which 
was being obtained from the small stockpile accumu¬ 
lated for war purposes. In an attempt to ease the 
cobalt situation, nickel-substituted stellites were 
tested with the unsatisfactory results already de¬ 
scribed in Section 19.6. In addition, a reconnaissance 
was made of a series of hot-hard alloys with less 
cobalt than Stellite No. 21 or with no cobalt. The 


melting points of these hot-hard alloys are all of the 
same order of magnitude as the stellites. 

19 8 2 Hot-Hard Reconnaissance 

Iron, Nickel, Chromium, and Cobalt Alloys 

The data for four groups of metals and alloys were 
examined to determine whether they would be resist¬ 
ant to melting and chemical attack by hot powder 
gases, resistant to deformation of the rifling by the 
swaging impact of the projectile, and sufficiently 
ductile to resist disintegration by brittle failure. 

Iron-Base Alloys. High-iron alloys (80% or more of 
iron) show poor resistance to powder-gas erosion. In 
the case of gun steel, performance is limited by a 
combination of chemical attack and flattening (swag¬ 
ing) of the rifling. With special steels or hot-hard, 
iron-base alloys or with ordinary steel hardened by 
nitriding or by special hardening techniques, chem¬ 
ical erosion by powder gases is still a serious limita¬ 
tion on performance unless the bore surface is pro¬ 
tected by an erosion-resistant plating or coating. 

Nickel-Base Alloys. High-nickel alloys (80% or 
more of nickel), with the exception of binary nickel- 
chromium alloys, show intergranular attack and dis¬ 
integration during firing. When the nickel-chromium 
alloys are hardened by alloying (for example, Inconel- 
type alloys) their resistance to powder-gas erosion is 
poor. 

Chromium-Base Alloys. Pure chromium and high- 
chromium alloys are very resistant to chemical attack 
and melting. However, most of these materials are 
too brittle for use as gun liners. Certain chromium- 
base alloys, discussed in Chapter 17, have a high 
melting point (1650 to 1700 C) and show excellent 
resistance to powder-gas erosion and very good hot¬ 
hardness. Some of these alloys have sufficient ductil¬ 
ity to withstand the shock of firing. 

Cobalt-Base Alloys. Pure cobalt and cobalt-base 
alloys (like the stellites) have good hot-hardness and 
some of them have sufficient ductility for satisfactory 
use as a gun liner. 

Hardened Iron-Nickel-Cobalt-Chromium 
Alloys 

Few data were available on the performance of 
hardened alloys that contain all four of the above 
metals. Many such alloys were known to have excel¬ 
lent high-temperature properties including hot-hard- 


CONFIDENTIAL 



ALLOYS CONTAINING LESS COBALT THAN MOST STELLITES 


407 


ness and hot-strength with small to moderate ductil¬ 
ity. The extent of their resistance to chemical attack 
powder gases was not known. In order to evaluate 
the possibilities of these alloys, a reconnaissance 
series of hardened iron-nickel-cobalt-chromium alloys 
was prepared and tested as liners in the caliber .50 
aircraft machine gun barrel. 80 - 81 The nominal com¬ 
position of the alloys subjected to firing tests is given 
in Table 5. They are specific examples of the types 
briefly described in Section 16.4.10, where their po¬ 
tentialities have been evaluated. The performance of 
none of these alloys was as good as that of invest¬ 
ment-cast Stellite No. 21. 

One alloy containing no cobalt (TEW) showed 
severe chemical attack by powder gases and shows 
no possibilities. Another alloy (Hastelloy C) contain¬ 
ing no cobalt was resistant to powder-gas erosion and 
shows considerable promise. The Hastelloys have al¬ 
ready been discussed in Section 16.4.9. The poor per¬ 
formance of some of the other alloys may have been 
caused by the poor condition of the specimen tested, 
owing either to inexperience in casting the particular 
alloy or to improper choice of hardness, which is 
dependent on the heat treatment and carbon con¬ 
tent. Thus the liner of N155 alloy that was tested was 
a poor sand-casting. A good, sound investment-cast¬ 
ing of this alloy with 0.25% carbon should be tested. 

One of the most promising of this group of hot-hard 
alloys is Refractaloy No. 70. 92 The first liner was bored 


Table 5. Nominal composition of hot-hard alloys (and 
of Stellites No. 21 and No. 27) used as caliber .50 liners 
in reconnaissance tests for erosion resistance. 


Nominal composition* 

Name Co Ni Fe Cr Mo W Other 


Stellite No. 21 f 

62 

2 

2 

28 

6 


C 0.25 

S 816f 

45 

20 

3 

20 

4 

4 

Cb 4 

Stellite No. 27 f 

32 

32 

2 

28 

5 


C 0.40 

Refractaloy No. 70§ 

30 

20 

15 

20 

8 

4 


N 155 f 

20 

20 

30 

20 

3 

3 

Cb 4 

MTB|| 

12 

30 

27 

20 

4 

4 

Ta 2 

Refractaloy No. 2§ 

15 

40 

13 

20 

5 

4 

Cb 3 

TEW|| 


30 

40 

20 

4 

4 

Ta 2 

Hastelloy “C”f 


58 

6 

15 

17 

4 


Hastelloy “A”f 
Hastelloy “B”f 


60 

20 


20 




66 

6 


28 




* All low carbon (usually 0.1%) unless otherwise stated; Mn and Si 
usually less than 1% each, 
t Haynes Stellite Company, 
t Allegheny Ludlum Steel Corporation. 

§ Westinghouse Research Laboratories. 

|| Special alloy similar to some studied in turbine blade research pro¬ 
gram. 169 

• 

from a forged and age-hardened bar. It showed very 
good resistance to powder-gas erosion but cracked 
during a second burst of 238 rounds after an initial 
burst of 350 rounds had been fired. Then investment- 
cast liners were prepared and tested. They were too 
soft and failed by swaging of the lands. The casting 
method should be perfected and heat treatment stud¬ 
ied to yield liners with optimum hot-hardness and 
ductility so that the erosion resistance can be utilized. 


CONFIDENTIAL 












Chapter 20 

ELECTROPLATING 

By William Blum a 


201 GENERAL PRINCIPLES 

T his chapter is confined to the application of 
electroplated coatings to gun bores to resist ero¬ 
sion. Only incidental references are made to the per¬ 
formance of the plated barrels, more details of which 
are given in Chapters 23 and 24. 

2011 Requirements 

Experience and research have shown (Chapter 16) 
that, to resist erosion, the material of the bore surface 
should be resistant to chemical attack by the powder 
gases, have a melting point of at least 1400 C, have 
high strength, hardness, and ductility at elevated 
temperatures, and have no abrupt volume changes 
with temperature. 

Because suitable metals are relatively scarce and 
may not be adaptable to the production of the entire 
gun tube, a “liner” or coating with the desired prop¬ 
erties is advantageous. Such a liner or coating may 
be produced mechanically, as for example, the molyb¬ 
denum and stellite liners (Chapters 18 and 19), by 
electrodeposition, by vaporization, by sputtering, or 
by chemical decomposition of a vaporized compound 
(Chapter 21). 

2012 Characteristics of 

Electrodeposited Coatings 

The advantages and limitations of electrodeposited 
coatings to protect gun bores may be summarized as 
follows. 

It is possible to vary the properties of the deposit 
by control of the conditions used in depositing a 
given metal. In the case of chromium plate the hard¬ 
ness may range from 400 to 1000 MVn. b Other prop¬ 
erties such as tensile strength vary over similar 
ranges. On heating to high temperatures the deposits 
of a given metal tend to anneal and to reach uniform 
properties for that metal. By co-depositing two or 

a National Bureau of Standards, U. S. Department of Com¬ 
merce. 

b Microvickers number: Vickers hardness determined with 
microhardness testing machine. 


more metals to form an alloy, it is possible to modify 
still further the properties of electrodeposits. 

The thickness of the electrodeposits can be con¬ 
trolled by regulating the current density and the 
period of deposition. The deposit follows closely the 
contour of the base metal, though the thickness is 
always less in the bottom of the grooves than on the 
top of the lands. The degree of this difference in 
thickness depends upon the throwing power of the 
plating bath employed. In chromium plating, the 
thickness of the deposit on the lands may be 30 per 
cent greater than that on the grooves since the chrom¬ 
ic acid plating bath has very poor throwing power. 
In other plating baths the difference is smaller. 

With chromium coatings up to 0.01 in. thick, the 
resultant changes in contour are not important ex¬ 
cept in artillery bores. The relatively higher and more 
sharply cornered lands in these bores may cause an 
undesirable build-up of the deposit on the land corners. 

With great care and, in certain cases, with special 
procedures, a degree of adhesion can be secured such 
that detachment of the coating removes part of the 
underlying metal. As no quantitative tests of adhe¬ 
sion are directly applicable to coatings in gun barrels, 
firing tests must be used as a criterion. 

2013 Availability of Electrodeposited Metals 

Only a few metals approach in their properties 
those desired in gun bores, that is high melting points 
and high hardness and tensile strength at elevated 
temperatures. The principal ones thus far considered 
for this purpose, as related in Section 16.1, are chro¬ 
mium, nickel, cobalt, molybdenum, tungsten and 
tantalum. Of these metals only the first three can be 
readily electrodeposited in a pure state from aqueous 
solutions. Available evidence indicates that it is not 
possible to electrodeposit pure tungsten or molyb¬ 
denum. 

Processes have been described for the electrodepo¬ 
sition of pure tungsten 463 but later published data 
indicate that deposits reported as tungsten were 
really alloys and contained at least small amounts of 
other metals, such as iron, nickel, or cobalt, that were 
present as impurities in the bath. By intentional ad- 


408 


CONFIDENTIAL 



CHROMIUM PLATING 


409 


ditions of such metals to baths of tungsten or molyb¬ 
denum it is possible to deposit corresponding alloys. 

There are a few reports in the literature on proces¬ 
ses for the electrodeposition of pure molybdenum but 
none of these processes yields satisfactory metallic 
coatings. A preliminary attempt aimed specifically 
at the preparation of a molybdenum-plated gun bar¬ 
rel was also unsuccessful because of poor adhesion of 
the molybdenum to steel. 80 

There are numerous patents on the electrodeposi¬ 
tion of tantalum from fused salt baths. The metal 
obtained is usually in the form of a fine powder, 
unsuitable to serve as a coating. More recently some 
attempts have been made without success to prepare 
an adherent coating from a fused bath consisting of 
potassium fluoride, potassium tantalum fluoride, and 
tantalum oxide at about 800 C. 95 

20 2 CHROMIUM PLATING 

Previous Experience 

More progress was made by Division 1 in the ap¬ 
plication of chromium than of other electroplates to 
gun barrels, because some of the properties of chro¬ 
mium approach those desired, and there had been 
previous experience in the use of chromium coatings 
to resist abrasion and corrosion. 171 Early attempts to 
apply it to gun barrels to resist erosion were not very 
successful. At the Washington Naval Gun Factory 
during the 1930’s chromium coatings less than 0.001 
in. thick were applied to the bores of large naval guns. 
These thin coatings did not greatly increase the use¬ 
ful life of these guns, 86 but they no doubt furnished 
some protection against corrosion, especially by sea 
water. Sporadic tests 185 by the Army Ordnance De¬ 
partment of gun barrels chromium plated commer¬ 
cially, and studies 311 at Frankford Arsenal on caliber 
.30 machine gun barrels, did not indicate any out¬ 
standing improvement in gun life. 

At the beginning of the Division 1 program, Service 
guns were carefully examined in order to determine 
in what manner the chromium plate had failed. 86 The 
conclusions drawn from these observations were 
strengthened by later studies of liners that had been 
fired in caliber .50 machine gun barrels in connection 
with the development of a technique for applying 
improved chromium plate to barrels of this caliber. 85 

The sequence of events in erosion of chromium- 
plated guns was found to be as follows. Microscopic 
cracks may be present in the plate as deposited (Fig¬ 


ure 1), and, because chromium is relatively brittle, 
additional cracks form during firing. The hot powder 
gases penetrate the cracks and react with the under¬ 
lying steel (Section 13.3) which has been thermally 
altered (Section 13.2). A pocket type of erosion, 
shown in Figure 1, takes place in the steel at the roots 
of the cracks in the chromium. The plate on the 
surface thus undermined, sometimes together with 
adherent steel, begins to be chipped and torn off by 
the projectile. This action is initiated most vigorously 
on the band slope and the adjacent edges of the lands. 



Figure 1 . Surface of chromium plate showing micro¬ 
scopic cracks in the metal as deposited: Unetched; 200X. 
(This figure has appeared as Figure 1 in NDRC Re¬ 
port A-414.) 

Such erosion progresses forward most rapidly on the 
tops of the lands, as illustrated in Figure 2. In the 
grooves, where contact with the projectile is slight, 
and on the chamber cylinder where no contact occurs, 
only minute chips of chromium are removed until 
corrosion along longitudinal cracks has greatly weak¬ 
ened the supporting steel. Swaging of the steel (Sec¬ 
tion 13.4.2), which may take place beneath the chro¬ 
mium plate, accelerates the cracking, undercutting, 
spalling, and mechanical removal of plate which has 
been cracked and pitted. 


CONFIDENTIAL 





410 


ELECTROPLATING 


Once the chromium plate is removed, erosion of a 
gun bore proceeds as in a nonplated gun. The initial 
cause of failure just described, however, is related to 
the mechanical weakness of chromium, which is re¬ 
sistant to thermal and chemical attack by the powder 
gases. As will be described later, the diminution of 
the mechanical forces can prolong the life of chro¬ 
mium-plated guns. The use of pre-engraved projec¬ 
tiles (Chapter 31) eliminates engraving stresses and 
minimizes bore friction, and the Fisa protector 
(Chapter 32) protects the chromium plate in the 
region of the origin of rifling. 



Figure 2. Cross section of eroded 5-in./25-cal., chro¬ 
mium-plated naval gun showing plate mostly in place 
in the groove but none on the land: Etched with picral; 
50X. (This figure has appeared as a portion of Figure 
17 in NDRC Report A-414.) 

20 2 2 Conditions of Deposition and 
Properties of Deposits 

Most commercial chromium plating is conducted 
from solutions containing from 250 to 400 g/1 (33 to 
55 oz/gal) of chromic acid and 2.5 to 4 g/1 of sulfuric 
add or an equivalent amount of another sulfate. The 
baths are usually operated at 40 to 60 C (104 to 
140 F) at current densities from 8 to 30 amp/dm 2 
(75 to 280 amp/ft 2 ). Under these conditions a bright, 
hard coating is produced with a cathode efficiency of 
about 15 per cent. For ornamental purposes, for ex¬ 


ample, over nickel on steel automobile parts, the 
chromium coatings are usually less than 0.00003 in. 
thick. For wear resistance on gauges and dies, coat¬ 
ings up to 0.005 in. thick may be used. 

Types of Chromium Plates 

Measurements made in the course of the investi¬ 
gation carried out by Division 1 show that the prop¬ 
erties of chromium deposits vary widely with the 
conditions of deposition. Two rather sharply defined 
types of chromium exist, to which have been applied 
the terms “HC” or “high contraction / 1 and “LC” or 
“low contraction.” Both types are included in Table 1. 
The HC deposits are of the type commonly used for 
decorative plating and for wear resistance. They are 
hard and brittle; they contain chromic oxide; and 
they contract when heated. In contrast the LC de¬ 
posits are much softer and less brittle (but not actu¬ 
ally ductile); they contain little oxide; and they 
contract only slightly when heated. Because of the 
lower contraction, the LC deposits are more nearly 
free from cracks, both as deposited and after having 
been heated. 

Plating Procedures for Different Base 
Materials 

No great difficulty is experienced in obtaining good 
adhesion of chromium to gun steel, if the customary 
methods of cleaning and pickling are employed, such 
as those listed under “Application to Caliber .50 
Barrels” in Section 20.2.3. 

Attempts to deposit chromium by these methods 
on stellite or similar alloys usually resulted in poor 
adhesion of the coatings, probably as a result of pas¬ 
sive films on the alloy surface. Plating on chromium 
and chromium alloys is often accomplished by initial 
deposition of nickel from an acidified nickel chloride 
solution. This method did not prove effective on 
stellite. 

Fairly good adhesion of chromium to stellite was 
obtained by scrubbing the surface first with mag¬ 
nesium oxide and water, followed by pumice and 
inhibited hydrochloric acid. A very low cathode cur¬ 
rent density (from 5 to 20 per cent of normal) was 
then applied for several minutes in the regular chro¬ 
mium plating bath, after which the current was grad¬ 
ually increased to normal during a period of 10 to 15 
min. Plating was then conducted as usual. This proc¬ 
ess represents cathodic pickling in the chromium 


CONFIDENTIAL 






CHROMIUM PLATING 


411 


Table 1. Composition and properties of electrolytic chromium deposited from a bath containing 250 g/1 Cr0 3 and 2.5 
g/1 S0 4 . 


Operating conditions 

Deposit 

Temp 

C 

C.D.* 

(amp/dm 2 ) 

C.E.f 

(%) 

Oxide 
content 
(% Cr 2 0 3 ) 

Contrac¬ 
tion on 
heating (%) 

As de¬ 
posited 

Hardness (MVn average values) 

Cooled after heating to: 
600 C 700 C 800 C 

1,000 C 

45 

25 

17 

1.4 

1.2 

850 

600 

500 

400 

200 

65 

30 

13 

0.7 

0.5 

800 





75 

50 

12 

0.2 

0.2 

500 

425 1 




85 

60 

10 

0.2 

0.2 

500 

350 

300 

200 


100 

80 




425 





20 § 

8 

40 

4.3 

1.8 

925 






* Current density, 
t Cathode efficiency, 
t Measured with Knoop Indentor. 

§ From U. S. Bureau of Mines trivalent chromium bath (see Section 20.4.1). 


bath. For example, for HC chromium, at 50 C, an 
initial current density of 1 amp/dm 2 was gradually 
increased to 20 amp/dm 2 ; while for LC chromium at 
85 C, the current density at the start was 15 amp/dm 2 
and was increased to 80 amp/dm 2 . 

One present limitation of plating on stellite liners 
is that it is difficult to remove stellite uniformly by 
electropolishing. Until this is accomplished, it will be 
necessary to machine liners oversize to allow for the 
thickness of chromium to be deposited. 

20 2 3 Applications to Gun Barrels 
Methods 

Both types of chromium were applied experiment¬ 
ally to gun barrels, chiefly to the caliber .50, by the 
National Bureau of Standards for Division 1. These 
experiments gave direct information regarding the 
performance of the coatings in small arms. 84 Some 


similar experiments were carried out at Springfield 
Armory. 86 - 318 

In plating the interior of a cylinder such as a gun 
barrel, a concentric inside anode is used to obtain 
approximately uniform distribution of the deposit. 
Typical fittings used to center and insulate the anode 
and to bring current to it and the cathode (the barrel) 
are shown in Figures 3 and 4. The longitudinal dis¬ 
tribution of the deposit is not usually uniform for 
the case where the only agitation is the liquid flow 
caused by upward pumping action of gas discharge, 
because it varies with the resistance of the anode and 
the surrounding solution as well as with other factors. 
By a suitable choice of the metal and the diameter of 
the anode it is possible to secure a wide range of 
distribution of deposits in a gun barrel of a given 
caliber. 

Thus it has been found 390 that with a steel anode of 
a given diameter, for example % in• in a caliber .50 
barrel, and with the anode connection at the top, the 



Figure 3. Fittings for electroplating caliber .50 aircraft machine gun barrel. (This figure has appeared as Figure 1 
in NDRC Report A-412.) 


CONFIDENTIAL 

















































412 


ELECTROPLATING 


deposit is thicker at the top than at the bottom. This 
behavior, caused by the high resistance of the steel 
anode, which reduces the current as it goes through 
the anode and thus lowers the current density and 
the thickness of the plate at the bottom, has been 
used to produce a tapered deposit and a desired 
“choke” near the muzzle of a caliber .50 aircraft 
machine-gun barrel. The advantages of the choked 
muzzle are discussed in Section 23.1.4. 



Figure 4. Rack for holding caliber .50 machine gun 
barrel during electroplating. (This figure has appeared 
as Figure 2 in NDRC Report A-412.) 


With a copper anode, which has a lower resistance 
than the steel anode, the longitudinal distribution is 
more nearly uniform, and the lower end of the barrel 
may even receive a thicker deposit than the upper. 
There are two factors which account for the latter 
type of uneven distribution. The temperature of the 
solution within the barrel is slightly higher at the top 
and this reduces the current efficiency. The gas 
bubbles formed by the evolution of hydrogen and 
oxygen accumulate at the top and increase the re¬ 
sistance of the solution, thus reducing the current at 
this end of the barrel. 

Both the steel and copper anodes are plated with 
lead, or preferably an alloy of lead with about 10% 
of tin, to prevent attack of the metal of the anode 
and to foster the reoxidation of trivalent chromium 
in the bath. At the cathode surface, some of the chro¬ 
mic acid (hexavalent chromium) is partially reduced 
and forms trivalent chromium, which in turn is re¬ 
oxidized at the anode until some equilibrium concen¬ 
tration is reached. 

The concentration of trivalent chromium affects 
the taper of the deposit and hence must be controlled. 
The cathode efficiency increases with an increase 
in trivalent chromium concentration in such a w^ay 


that the increase in current efficiency at low current 
densities is greater than that at higher current den¬ 
sities. Therefore with a steel anode, which produces a 
lower current density at the bottom than at the top 
of the barrel, the presence of trivalent chromium 
results in a higher current efficiency and hence rela¬ 
tively higher plating rate at the bottom than at the 
top of the barrel. 

The deposit taper, that is, the ratio of the thickness 
at the top to that at the bottom of the bore, therefore 
decreases with increase in trivalent chromium con¬ 
centration. In the caliber .50 machine-gun barrel, 
plated with a % 5 -in. steel anode, this ratio is approx¬ 
imately 3.0 when the trivalent chromium concentra¬ 
tion is nearly zero, and is 1.5 when the trivalent 
chromium concentration is 15 g/1. The concentration 
of trivalent chromium can be reduced by electrolysis 
of the solution in a separate operation with large 
anodes at a low T current density and with small dummy 
cathodes, preferably surrounded by a porous pot. 

In most of the plating experiments at the National 
Bureau of Standards the gun barrel with the assem¬ 
bled anode and fittings w r as immersed in a tank con¬ 
taining the plating solution that w^as maintained at 
the specified temperature. Holes in the fittings per¬ 
mitted the circulation of the plating solution, induced 
by a rising stream of gas, consisting of oxygen from the 
anode and hydrogen from the cathode. 

Much of the exploratory plating was done on short 
steel breech liners, subsequently inserted into the 
heavy caliber .50 barrels used in rapid-fire tests at the 
Geophysical Laboratory 81 at normal velocities and in 
other tests with the hypervelocity erosion-testing gun 
(Section 11.2.1) at the Franklin Institute. 76 The LC 
chromium gave approximately 23 per cent greater 
increase in the life of the barrel fired at normal veloci¬ 
ties than did the HC chromium, but conversely, in 
the aircraft barrels, the HC was superior. This differ¬ 
ence illustrates the variation in the nature and degree 
of erosion in barrels fired under different schedules or 
having different construction and operation. 

The production of a tapered deposit by means of a 
steel anode as described above, led to a muzzle choke 
that greatly increased the accuracy life of the aircraft 
barrels as detailed in Chapter 23. The HC chromium 
prevented erosion at the muzzle and hence preserved 
the restricted diameter. 

If more than 0.001 in. of chromium is to be applied, 
especially at the origin of rifling, it is necessary to 
remove sufficient steel to make room for the chro¬ 
mium. In production of new T barrels for plating, this 


CONFIDENTIAL 





























CHROMIUM PLATING 


might be accomplished by reaming and rifling to 
oversize diameters. Before plating barrels of standard 
dimensions steel was removed by “electropolishing” 
(see outline in Section 25.2.1). This process 0 was 
developed in recent years as a means of producing a 
bright surface on metals. It depends on the anodic 
solution of metal in a suitable electrolyte. d 

The solution finally adopted for gun barrels con¬ 
sists of equal volumes of 96% sulfuric acid and 75% 
phosphoric acid. The polishing is conducted at 43 C 
(110 F) and an anodic current density of 27 amp/dm 2 
(250 amp/ft 2 ). The barrel is made anodic, and a 
copper or steel rod, coated with lead-tin, is used as 
the cathode. The fittings are similar to those in Fig¬ 
ures 1 and 2. In addition to serving as a “taking off 
tool,” electropolishing may be beneficial in improving 
the surface condition of the steel and fostering good 
adhesion of the deposits. Electropolishing is a neces¬ 
sary step in the preparation of nitrided steel barrels, 
from which the outer brittle layer, approximately 
0.001 in. thick, must be removed before plating in 
order to obtain good adhesion of the deposit. 

Application to Caliber .50 Barrels 

Nitrided Barrels. Firing tests of caliber .50 aircraft 
barrels that had been plated with chromium directly 
on gun steel showed that failure of the chromium 
resulted in part from the swaging of the underlying 
steel as described in Section 13.4.2. It was then de¬ 
cided to harden the bore surface by nitriding to give 
a better foundation for the chromium deposit. This 
resulted in a marked improvement that led to the 
adoption of the nitrided chromium plated aircraft 
barrel by the War Department, as described in Chap¬ 
ter 23. 

On the basis of firing tests by the Geophysical 
Laboratory of several hundred nitrided barrels plated 
at the National Bureau of Standards, specifications 
were drawn up. These were employed in a pilot plant 
at the Doehler-Jarvis Corporation in Grand Rapids, 
Michigan, and minor changes were made before pro¬ 
duction of plated barrels, described in Chapter 25, 
was undertaken in a larger unit by that company. 
The bullet seat must be reamed out to 0.515 + 0.002 
in. before nitriding to lead to the specified final di- 


c An extensive bibliography on the subject in reference to 
types of electrolytes is listed in a paper by Zmeskal. 470 

d Valuable information was obtained informally from C. L. 
Faust of Battelle Memorial Institute. 


413 


mension. If this is not done, constriction of the bullet 
seat will result from swaging during firing and will 
cause jamming of the bullets. 

The exact methods and conditions for preparing 
and plating the barrels w^ere not specified, but the 
following procedure w r as recommended and generally 
followed. (The steps are described in more detail in 
Section 25.2.1.) 

1. Degrease. 

2. Decopper (if necessary), rinse, dry. 

3. Gauge. 

4. Electropolish to remove 0.001 to 0.002 in. of 
metal with either end up. 

5. Rinse. 

6. Scrub with pumice and inhibited hydrochloric 
acid or with pumice and an alkaline cleaner. 

7. Dry. 

8. Oil (if to be stored). 

9. Gauge. 

10. Degrease. 

11. Scrub with pumice and acid. 

12. Rinse. 

13. Dry with patch. 

14. Etch anodically in chromic acid for 5 min. 

15. Plate with chromium, with muzzle end up, 
using a %-in. steel anode, at 50 C (122 F) and 20 
amp/dm 2 (190 amp/ft 2 ), for the calculated time 
(about 4 hr). 

16. Rinse. 

17. Dry. 

18. Gauge. 

19. Oil for storage. 

Stellite Liner Barrels. When it was realized, as 
described in Section 24.1.1, that aircraft barrels pro¬ 
vided with a stellite liner and chromium plate beyond 
the liner w ould yield better service than the nitrided 
plated barrels, development of this “combination” 
barrel w r as undertaken and a specification was pre¬ 
pared based on plating and firing tests. The essential 
requirements w r ere as given in Table 2. 


Table 2. Specifications of finished dimensions for 
chromium plate in caliber .50, 36-in. barrel with 9-in. 
stellite liner. 


Distance from muzzle 
(in.) 

Diameter 

(in.) 

0.5 or 1.5 

0.4920 + 0.0045 

10 

0.4950 + 0.0035 

15 

0.4955 + 0.0035 

23 

0.4990 + 0.0020 

Thickness of deposit at 23.5 in. from muzzle: 0.0017 ± 0.0007 in. 


CONFIDENTIAL 









414 


ELECTROPLATING 


Several methods were tried for plating liner-barrels. 
From the plating standpoint, deposition of chromium 
before liner insertion is preferable; but from the 
standpoint of production efficiency, plating after in¬ 
serting the liner is better. The latter method proved 
difficult because the solutions tended to enter and be 
retained in the crack at the forward end of the liner 
and to cause etching of the liner. In the production 
plating of liner-barrels by the Doehler-Jarvis Corpor¬ 
ation (Chapter 25), a dummy steel liner was inserted 
in the barrel, which was then electropolished and 
plated under the same conditions as full-length bar¬ 
rels. The dummy liner was counterbored at the for¬ 
ward end to avoid deposition of a sharp edge of chro¬ 
mium there. Further study is required to define the 
best method of plating liner-barrels. 

Heavy Barrels. Onty a small number of the 45-inch 
caliber .50 barrels was plated experimentally. Just as 
with the plated liners inserted in these barrels, a full- 
length deposit of LC chromium, about 0.006 in. thick 
at the origin, yielded the best results. It was not 
possible, however, to secure the desired longitudi¬ 
nally uniform deposit of LC chromium in these bar¬ 
rels by any variations in the anode size or composition. 

Use of a moving anode, regularly employed in 
chromium plating large naval guns with HC chro¬ 
mium, yielded LC deposits with poor adhesion. Ap¬ 
parently the exposure of part of the bore to the 
chromic acid bath at the higher temperature caused 
excessive passivity of the steel so exposed. 

The best LC deposits were obtained by “pump 
plating.” With a well-conducting anode, such as of 
copper, silver, or aluminum, any taper caused by the 
anode resistance is negligible. Any reverse taper, 
caused by a difference in the bubble concentration 
and temperature at the two ends, is avoided by pump¬ 
ing the solution through the bore at a high rate of 
flow. This procedure reduces the difference in bubble 
concentration and temperature at the two ends and 
the higher pressure involved also reduces the bubble 
volume. The current efficiency is normal at a flow 
rate of 100 ft/min (30 m/min). By varying the flow 
rate, the longitudinal distribution of the deposit can 
be controlled to give either a straight or a choked 
bore. 

To avoid passivating the steel by the hot chromic 
acid during the change from the 50 C etching solution 
to the 85 C plating bath, the bore was plated, after 
etching, with about 0.0002 inch of HC chromium, 
followed by the LC chromium without current inter¬ 
ruption. 


Miscellaneous Applications 

Firing tests on a small number of caliber .30 barrels 
plated with HC chromium yielded results parallel to 
those with the caliber .50 aircraft barrels. A deposit 
about 0.0025 in. thick at the origin and tapered to 
effect a muzzle choke of about 0.006 in. in diameter 
produced a decided increase in life under a severe 
firing schedule. 81 A few experiments with pump-plat¬ 
ing in caliber .30 barrels produced an LC deposit with 
satisfactory dimensions, but firing tests have not yet 
been made. 

A number of caliber .60 barrels were plated, princi¬ 
pally with HC chromium. Firing tests were most 
promising in barrels plated with from 0.005 to 0.010 
in. of HC chromium for several inches from the 
breech and with only 0.001 to 0.002 in. of plate from 
there to the muzzle, which was not “choked” in these 
barrels. The life of this caliber .60 barrel is at least 
doubled by this type of deposit. More tests are 
planned, including barrels plated with LC chromium. 

In one test of a 20-mm barrel plated with 0.01 in. 
of HC chromium, the barrel life was approximately 
doubled. The most favorable thickness and distribu¬ 
tion of chromium for this gun barrel have not yet 
been defined. 

Two hypervelocity 37-mm gun barrels, T47, de¬ 
scribed in Section 31.7, and the 37-mm gun barrel 
used for tests of the Fisa protector, described in Sec¬ 
tion 32.5, were plated at the Washington Naval Gun 
Factory with the cooperation of the National Bureau 
of Standards. Not enough data were obtained to 
warrant recommendations for the type or thickness 
of chromium. 

Work done at Battelle Memorial Institute 271 ’ 272 ’ 276 
on 37-mm barrels indicates that there is a continuous 
improvement in the performance of such barrels 
plated with HC chromium with increase in deposit 
thickness up to 0.010 in. 

Parts of a 57/40-mm tapered-bore gun, a 4.7-in. re¬ 
coilless mortar and some 75-mm recoilless rifles were 
plated at the National Bureau of Standards and tested 
by different agencies. In general the chromium coat¬ 
ing was beneficial, but the results were not sufficiently 
numerous or consistent to warrant definite recom¬ 
mendations. 

As previously noted, most of the plating heretofore 
applied to large guns was less than 0.001 in. thick. 
During the past few years several large guns were 
plated with up to 0.006 in. of HC chromium by the 
Washington Naval Gun Factory for experimental 


CONFIDENTIAL 



NICKEL AND COBALT PLATING 


415 


firing by the Army Ordnance Department. In one 
test , 216 a plated 155-mm barrel showed excessive muz¬ 
zle erosion, and the chromium made no significant 
improvement. The data on this single test are incon¬ 
clusive. 

While the work described in the foregoing para¬ 
graphs was being carried out in the United States, 
efforts were also being made in Great Britain to im¬ 
prove chromium electroplates for application to gun 
barrels . 380 - 385 ’ 386 Attention was concentrated especial¬ 
ly on plating caliber .50 machine-gun barrels , 389 - 390 ’ 423 
with results that compared favorably with those 
achieved by Division 1 . German endeavors in this 
field did not result in much improvement . 306 ’ 391 

20 2 4 Duplex Coatings 

The important part played by cracks in the failure 
of a chromium electroplate on the bore surface of a 
gun (Section 20.2.1) led to the suggestion that im¬ 
proved performance might be obtained by the use of 
an undercoat of some other, more ductile material. 
Duplex coatings consisting of chromium deposited on 
top of various materials were tested in the caliber .50 
erosion-testing gun (Section 11.2.1). 

Similar behavior was shown by 1 -mil plates of 
copper and of nickel beneath a 1 -mil chromium plate. 
The undercoat became plastic, whereupon the en¬ 
graving stresses rubbed the chromium plate off the 
lands. The copper plate successfully sealed the bore 
surface from attack by the powder gases, whereas the 
nickel plate cracked and permitted the powder gases 
to reach the steel. A combination of a 1-mil plate of 
nickel on a 1 -mil plate of copper under a 1 -mil plate 
of chromium behaved in the same way as the nickel 
undercoat. 

In experiments with the erosion-testing gun, de¬ 
scribed in Section 31.5, it was found that a thickness 
of chromium of at least 6 mils on the lands was neces¬ 
sary to prevent thermal alteration of the underlying 
gun steel . 77 A plate of this thickness would presum¬ 
ably reduce the softening of copper in a duplex plate. 
The trial of this combination had been considered at 
one time, but was not carried out because of other 
more urgent tests. It might be worthwhile to make a 
systematic study of the effect of varying the propor¬ 
tions of the copper and chromium layers in a duplex 
coating. 

An even more promising type of undercoat for 
chromium plate, is one of cobalt or of a cobalt-tung¬ 
sten alloy plate. Three tests were made of chromium- 


cobalt duplex coatings in the caliber .50 erosion-test¬ 
ing gun. There were 3 mils of chromium on 7 mils of 
cobalt. Excellent protection of the gun steel surface 
against powder gas erosion was reported , 76 - 77 but the 
cobalt undercoat was swaged, and eventually the 
plate was removed from the lands for a short distance 
ahead of the origin of rifling. 

Cobalt-tungsten alloy plates of the sort described 
in Section 20.4.2 are considerably harder than pure 
cobalt. Therefore, duplex plates were prepared with a 
7-mil undercoat of one of these alloys (containing 
10 % tungsten) under 2 mils of chromium. These du¬ 
plex plates did not last as long as those containing 
pure cobalt undercoats; but this result should not be 
construed as final. The development of these alloy 
plates is still in a very early stage, and hence it is not 
certain that the particular ones tested represented 
the best results possible. There is still the hope that 
it may be possible to develop one of these alloy plates 
so that it will have just the right combination of 
properties to provide the perfect undercoat for chro¬ 
mium plate for the surface of a gun bore. 

20 3 NICKEL AND COBALT PLATING 

It is possible to electrodeposit nickel and cobalt 
that have a range of hardness from about 100 to 
400 MVn. The nickel and cobalt which have been 
deposited in gun bores have approximately the same 
hardness as gun steel (280 to 320 MVn). Therefore 
no appreciable increase in resistance to abrasion or to 
swaging should be expected through the use of a 
surface layer of nifckel or cobalt. This assumption is 
borne out for nickel by the experience of the British 
Armament Research Department , 367,380,386 which 
made extensive tests on the use of nickel deposition 
to salvage worn-out artillery barrels . 379,382,383 By 
carefully defined technic they secured good adhesion 
of heavy nickel deposits which, after having been 
machined and rifled, yielded about the same service 
as new gun-steel barrels. 

Cobalt may be deposited from baths similar to 
those used for nickel plating. Very satisfactory depos¬ 
its of cobalt were obtained from a simple solution con¬ 
taining 400 to 500 g /1 of cobalt chloride, CoCl 2 * 0 H 2 O. 
The pH was kept between 3 and 4.5. At room tem¬ 
perature, a current density of 2 to 5 amp/dm 2 was 
used. 

Good adhesion of the cobalt coatings on steel was 
obtained by first etching the steel anodically in 70% 
H 2 SO 4 for 2 min at 25 amp/dm 2 . A “strike” coating 


CONFIDENTIAL 



416 


ELECTROPLATING 


of cobalt was applied for 3 min at 20 amp/dm 2 in a 
solution containing about 100 g/1 of cobalt chloride, 
kept at a pH of 0.5 to 0.7 with HC1. Plating was then 
conducted from the stronger solution above described. 

Electroplated cobalt is usually harder than the 
electroplated nickel commonly used. Nickel plate 
softens considerably after annealing at 800 C whereas 
cobalt retains most of its initial hardness. 

Cobalt possesses better erosion resistance than 
nickel. The firing tests described in Sections 16.3.1 
and 16.4.9 have shown that nickel is likely to erode as 
a result of intergranular corrosion. 

A liner was plated with 0.005 in. of cobalt and 
tested in the caliber .50 erosion-testing gun. (Section 
11.2.1). The adherence of the cobalt was good, but 
its resistance to melting and to swaging was less than 
that of chromium. 76 

26 4 ALLOY DEPOSITION 

A study of the plating of alloys is warranted be¬ 
cause, as above noted, such metals as tungsten and 
molybdenum may be depositable only as alloys; 
moreover, by the co-deposition of two metals in con¬ 
trolled proportions, certain desired properties may be 
obtained. While it is possible to co-deposit three or 
possibly more metals, the definition and control of 
such processes are much more complicated than for 
binary alloys. 

20 4 1 Chromium Alloys 

Because chromium has been found to improve the 
performance of gun barrels, efforts were made to de¬ 
posit alloys of chromium that might be less brittle 
and have less tendency to crack than pure chromium. 
It is possible to introduce such metals as iron, nickel 
and cobalt in the form of dichromates into the reg¬ 
ular chromic acid baths. 460 It is difficult, however, to 
co-deposit more than a few tenths of 1 per cent of 
nickel or cobalt with the chromium and more than a 
few per cent of iron 461 ’ 462 under the conditions for 
production of either HC or LC chromium. The result¬ 
ant deposits have essentially the same properties as 
those of pure chromium deposited under the same 
conditions. 

Relatively soft chromium deposits at 85 C and 

21 amp/dm 2 (200 amp/ft 2 ) were produced in England 
at first with a low cathode efficiency. The efficiency 
was then increased by adding 25 g/1 of iron or 20 g/1 
of trivalent chromium to the bath. It is possible to 


add as much as 45 g/1 of iron, but such baths are 
unstable. Besides increasing the cathode efficiency, 
addition of 25 g/1 or more of iron widens the permis¬ 
sible sulfate range. 384,471 

Thick deposits from the iron-chromic acid baths at 
85 C are sounder, that is, less likely to spall or crack, 
than similar deposits from the chromic acid bath. It 
is not possible to state whether these differences in 
behavior result directly from the small iron content 
of the deposit. 

The slight difference in the properties of the plate 
from the alloy dichromate bath as compared with 
that from the chromic acid bath under similar con¬ 
ditions did not justify the use of such plating baths 
in the investigation at the National Bureau of Stan¬ 
dards, since they are difficult to prepare and some¬ 
what unstable at high concentrations. It was found 
that the current efficiency of the chromic acid bath 
can be more readily increased, particularly at low 
current densities, either by diluting the bath or by 
adding hydrofluoric acid instead of sulfuric acid. 
Neither of these baths has been used in the plating of 
gun barrels. If the physical and mechanical proper¬ 
ties of the deposits from these baths prove to be as 
good as those from the regular chromium baths, they 
should be tried in gun barrels. 

Alloys of chromium with as much as 1% of tung¬ 
sten or molybdenum were obtained from chromic 
acid baths containing added tungstate or molybdate 
and fluorides or phosphates. The cathode efficiencies 
were very low. 

Efforts were made to co-deposit other metals with 
chromium from baths containing chromic or chro- 
mous salts. 459 Recently the U.S. Bureau of Mines 469,522 
developed a process of recovering chromium from its 
ores by electrolysis of a bath containing chromic and 
chromous salts and sodium sulfate at a pH of 1.8. An 
insoluble anode was surrounded by a diaphragm. 
Efficiencies as high as 40 per cent were obtained, but 
the chromium deposits were dark and brittle and 
contained as much as 4% of chromic oxide. (See 
Table 1.) It was not found possible to obtain any 
promising alloy deposits from baths of this type or 
from alkaline chromium baths. 

Deposits consisting of cobalt with only 1% of 
chromium were obtained from baths containing 100 
g/1 of cobalt as sulfate, 50 g/1 of chromium as chromic 
sulfate, and hydroxyacetic acid. The hardness of the 
deposit was somewhat higher than that of pure 
cobalt. 

From a bath containing tungstate, a ternary 


CONFIDENTIAL 



CONCLUSIONS AND RECOMMENDATIONS 


417 


deposit with about 1% of chromium, 5% of tung¬ 
sten, and the balance cobalt, was obtained. Further 
work is required to determine whether deposits of 
this type are practicable and useful for this purpose. 

Alloys Containing Tungsten 

Between 1930 and 1940 the Tungsten Electrode¬ 
posit Corporation patented acidified baths contain¬ 
ing fluorides for the deposition of alloys of tungsten 
with nickel or other metals. 527 Other investigators at 
the University of Wisconsin published methods for 
depositing tungsten alloys from both acid and alka¬ 
line baths. 464 ' 465 466 - 467 ' 468 There are Russian publica¬ 
tions on deposition of these alloys from ammoniacal 
baths. 472 - 473 - 474 

Experiments 84 at the National Bureau of Stan¬ 
dards did not find it possible to produce dense 
coherent alloy deposits from any of the baths de¬ 
scribed above. It was then found possible to obtain 
satisfactory deposits from ammoniacal solutions con¬ 
taining salts of hydroxy-organic acids. 

A typical bath for depositing alloys of cobalt and 
tungsten contains 25 g/1 of cobalt (as sulfate or chlo¬ 
ride), 10 g/1 of tungsten (as sodium tungstate), 400 
g/1 of Rochelle salt (sodium potassium tartrate) and 50 
g/1 of ammonium chloride. The pH is adjusted with 
ammonium hydroxide to 8.5, and the bath is operated 
at 90 to 100 C and at 1 to 5 amp/dm 2 . 

Deposits containing from 10 to 35% of tungsten 
and the balance cobalt, nickel, or iron were obtained 
from baths of this type. Deposits up to 0.05 in. thick 
were smooth and strong, but brittle. The most prom¬ 
ising alloys were the cobalt-tungsten ones, the hard¬ 
ness of which, as deposited, ranges from 500 to 700 
MVn. The MVn of iron-tungsten alloys with 50% 
tungsten is from 700 to 1,000. 

When these alloys are heated to 600 C for an hour 
and cooled, their hardness increases by as much as 
100 MVn for cobalt-tungsten or nickel-tungsten, and 
200 for iron-tungsten. Heating these alloys to 900 C 
or higher causes them to soften permanently. The 
cobalt-tungsten alloys have hot-hardness values that 
are higher than those of chromium, and unlike chro¬ 
mium, they retain their hardness on cooling. 

It is difficult to obtain good adhesion of the cobalt- 
tungsten alloy directly to gun steel. The procedure 
finally adopted was to plate a thin layer of cobalt on 
the steel, treat the cobalt in the alloy bath with an 
alternating current, and then deposit the alloy. De¬ 
posits thus applied to liners had fairly good adherence 


but in firing tests they showed some flaking near the 
origin. On heating the plated liner to 900 or 1000 C, 
the adherence of the alloy was improved, and the 
deposit was rendered more ductile. This heating, 
however, tends to crack the alloy layer and to soften 
the gun steel. 

Liners plated with nickel-tungsten and fired in the 
caliber .50 erosion-testing gun suffered severe gas 
erosion. 76 - 77 On the other hand, liners plated with 
cobalt-tungsten showed 76 better resistance to erosion 
than gun steel but were inferior to those plated with 
pure cobalt or chromium. The adhesion of the heat- 
treated alloy was better than that of chromium. 
There was less swaging of the lands with the cobalt- 
tungsten than with cobalt. Efforts are being made to 
obtain satisfactory adhesion of chromium to the 
cobalt-tungsten alloy, in order to make possible du¬ 
plex plates of the sort described in Section 20.2.4. 

Much further work, including possible modifica¬ 
tion of the gun steel, is required to realize the full 
possibilities for alloys of tungsten with cobalt or 
other metals. For example, it may be possible to 
produce by electroforming on a suitable mold, liners 
having an alloy surface and a body of another metal. 
Some exploratory trials gave promising results. 84 

20 4 3 Alloys Containing Molybdenum 

Much less progress was made in the deposition of 
alloys of molybdenum. The most promising bath 
contains 1,000 g/1 of potassium carbonate, 15 g/1 of 
cobalt as sulfate or chloride, and 100 g/1 of sodium 
molybdate. At a pH of 11, at 100 C and 1 to 5 
amp/dm 2 , deposits containing up to 35% of molyb¬ 
denum are produced. They are as hard as the tung¬ 
sten alloy deposits and harden on heating but are not 
as strong as the tungsten alloys. 

20 5 CONCLUSIONS AND 

RECOMMENDATIONS 

Chromium Deposits 

The successful application of HC chromium to 
caliber .50 aircraft barrels, either previously nitrided 
(Chapter 23) or provided with a stellite liner (Chap¬ 
ter 24), illustrates the need for study of each type of 
w r eapon and condition of service. More experience is 
required in the plating and firing of different small- 
caliber guns, including their examination to deter¬ 
mine the behavior of each coating at the breech and 


CONFIDENTIAL 




418 


ELECTROPLATING 


muzzle, and the causes of failure. Such information 
may be valuable not alone in improving the perform¬ 
ance of small arms, but also in indicating the most 
promising materials for artillery, on which the firing 
tests are necessarily more restricted. Such correlation 
depends on more complete data on the temperatures, 
pressures, and erosive conditions involved in each 
weapon. Such studies should include different types 
and thicknesses of chromium, applied by methods 
that may change the adherence, distribution or prop¬ 
erties of the deposits. 

20 5 2 Duplex Coatings 

The use of composite metal coatings, for example, 
of cobalt or alloys followed by chromium, warrants 
further study. 

20 5 3 Alloy Deposits 

The studies thus far conducted at the National 
Bureau of Standards and elsewhere on the deposition 
of alloys of tungsten or molybdenum with iron, 
nickel, or cobalt, are exploratory and valuable in 
showing that such alloys can be deposited in a dense 
form and that they possess a wide variety of proper¬ 
ties. Much more research is required to define and 
control favorable conditions for their application to 
gun bores. They are more likely to prove useful in 
large than in small guns, because, in the latter, liners 


of stellite or molybdenum can be inserted. Alloy 
liners may possibly be produced by electroforming. 

20.5.4 Properties of Electrodeposits 

Much intensive study is required on the properties 
of electrodeposited metals and alloys, including the 
hardness at room and elevated temperatures, and 
after having been heated; tensile strength, ductility, 
elastic properties, coefficient of expansion (including 
permanent expansion or contraction) and resistance 
to chemical attack by air, water, or powder gases. 
The existing data on well-known metals, such as 
chromium, nickel, and cobalt, are fragmentary and in 
some cases contradictory. Acquisition of more reli¬ 
able data should lead to more successful applications 
of deposited metals and alloys for both military and 
industrial purposes. 

20 5 5 Properties of Steel 

Further studies on the properties of the steels used 
in gun barrels are necessary in order to permit the 
full possibilities of electrodeposited coatings or liners 
to be realized, as is brought out in Section 24.5. The 
composition and properties of the steel may affect the 
adhesion and performance of the coating and may 
determine what steps, such as heat treatment, may 
be used to improve the properties or adhesion of the 
coatings. 


CONFIDENTIAL 



Chapter 21 

VAPOR-PHASE PLATING OF MOLYBDENUM, TUNGSTEN, 

AND CHROMIUM" 

By Charlotte A. Marsh h 


211 INTRODUCTION 

21,1,1 New Method of Plating Refractory 
Metals 

OLYBDENUM, TUNGSTEN, AND CHROMIUM have 

highly desirable erosion-resistant properties. 
This was demonstrated by tests described in Chap¬ 
ters 16, 17, and 18. The supply of these metals, which 
is limited in normal times, is especially short in war 
time because of other large demands of high priority. 
For this reason it was expedient to consider their 
application to gun-bore surfaces as thin plates. More¬ 
over, massive molybdenum requires a large amount 
of mechanical working to provide a suitable liner 
material, necessitating the elaborate and expensive 
equipment described in Chapter 18; massive tungsten 
is virtually unworkable; and massive chromium is 
brittle, which was learned from the experiments de¬ 
scribed in Chapter 17. 

Thin coats of metals are ordinarily applied by elec¬ 
troplating. Experiments in applying electroplates of 
these metals are described in Chapter 20. Chromium 
has proved very successful when applied to steel bore 
surfaces, but satisfactory electroplates of molyb¬ 
denum and tungsten have not yet been prepared. 
Chromium electroplates, however, do not adhere to 
the surface of stellite or alloys of this type. These 
alloys, as stated in Chapter 19, have comparatively 
low melting points, hence their use as erosion resist¬ 
ant liners is somewhat limited by the temperature 
conditions to be met in firing. The application of a 
material of higher melting point, such as molybdenum 
or chromium, to stellite liners may extend their use¬ 
fulness. This use necessitates the development of a 
method for applying adherent coatings of these 
metals to stellite. 


a This chapter has been based entirely on four Division 1 
formal reports, 73 ’ 74 ’ 93 ’ 94 to which reference is made for further 
details. 

b Geophysical Laboratory, Carnegie Institution of Washing¬ 
ton. (Present address: U. S. Geological Survey, Washington, 
D.C.) 


It was thought both possible and practical to pre¬ 
pare adherent (to steel or stellite) thin plates of 
molybdenum, tungsten, and chromium by thermal 
decomposition of the vapors of their respective 
carbonyls. To this process, which appears to be novel, 
the descriptive phrases “vapor-phase plating’ ’ and 
“pyrolytic plating” have been applied. The former is 
suitable because the metal plate is deposited by de¬ 
composition of a vapor, not from solution; the latter 
because the vapor is decomposed by heat. 

21,1,2 Carbonyl Chemistry 447 * 448 

Certain of the so-called “heavy metals,” notably 
iron, nickel, cobalt, tungsten, molybdenum, chromi¬ 
um, and the platinum metals, have the property of 
combining readily with carbon monoxide gas under 
appropriate conditions to form reasonably stable 
chemical compounds known as carbonyls. These have 
the general chemical formula M(CO)*, where M is 
the metal and x is four or greater. The reaction be¬ 
tween the metal and carbon monoxide takes place at 
an appreciable rate only under limited ranges of tem¬ 
perature and pressure, and these vary from metal to 
metal. Except for iron and nickel, these conditions are 
not readily obtained, and the other carbonyls are 
usually prepared by reaction between carbon monox¬ 
ide and an appropriate compound of the metal. 

The carbonyls, when once formed, are stable only 
within definite, limited ranges of temperature and 
pressure. When the temperature at a given pressure 
is raised beyond a specific point, the carbonyl de¬ 
composes to reform the metal and carbon monoxide. 
A secondary reaction may also occur in which carbon 
and carbon dioxide are formed from the monoxide. 
The carbon thus released may be deposited on an ad¬ 
joining surface together with the metal resulting from 
the decomposition, or it may react with it to form one 
or more carbides. This secondary reaction is of con¬ 
siderable importance, as will be seen in due course. 
The amount of carbon or carbide formed may be 
varied by suitable alteration of the experimental con¬ 
ditions. 



CONFIDENTIAL 


419 



420 


PLATING OF MOLYBDENUM, TUNGSTEN, AND CHROMIUM 


The formation and decomposition of nickel car¬ 
bonyl is the basis of the well-known Mond process for 
the manufacture of pure nickel. The formation and 
decomposition of iron carbonyl 6 is employed on a lab¬ 
oratory scale for the preparation of small samples of 
highly pure iron for scientific purposes, and has lately 
been used for the preparation of pure iron in quanti¬ 
ties of several pounds at a time for the manufacture 
of iron articles by powder metallurgy. 

211,3 Summary of Recent Progress 

The carbonyls of molybdenum, tungsten, and 
chromium have been known for some years. Surveys 
of the literature 73,93,94 revealed that they have only 
been prepared on a laboratory scale and that no 
attempts have been made hitherto to utilize either 
the metals or the carbonyls prepared by their de¬ 
composition. Improved methods for the preparation 
of molybdenum and chromium carbonyls have now 
been developed as described in Sections 21.2 and 21.5, 
respectively. The considerable work carried out on 
the vapor-phase plating of molybdenum is summa¬ 
rized in Section 21.3. Only a preliminary investiga¬ 
tion on the preparation of tungsten carbonyl, the 
preparation of plates therefrom, and the properties of 
these plates has been made, since molybdenum is 
more readily available. The use of cobalt plate's pre¬ 
pared by this method has not been taken up because 
of the short supply of cobalt, and its higher priority 
use as an essential constituent of stellite and similar 
alloys. 

None of the plates prepared was satisfactory for 
the bore surface of a gun, because the high tempera¬ 
ture produced during firing weakened the bond be¬ 
tween the plate and the underlying material. 

The hardness of the plates that are obtained by the 
method of vapor-phase plating that has been devel¬ 
oped can be varied at will. Hence, with a little fur¬ 
ther investigation, they may have practical application 
for the hard-surfacing of metals for use where resist¬ 
ance to mechanical wear without severe mechanical 
shock is involved. If considerably thicker plates of 
good adherence can be prepared, there is reasonable 
prospect of success in obtaining resistance to high 
temperatures. 


c A detailed study of the reaction between carbon monoxide 
and iron was carried out 19 preliminary to investigations of the 
possible importance of the formation of iron carbonyl in gun 
erosion. 62 ’ 63 This work is summarized in Section 14.3. 


212 PREPARATION OF MOLYBDENUM 
CARBONYL 94 

21.2.1 Formation of Molybdenum Carbonyl 

The method which was developed for the prepara¬ 
tion of molybdenum carbonyl in amounts adequate 
for this program utilized the reaction of carbon mon¬ 
oxide with a reactive form of molybdenum at high 
pressures. The use of metallic molybdenum as such is 
not practical, for, no matter how finely divided the 
metal may be, it reacts too slowly with carbon mon¬ 
oxide. The reason for this is obscure, but is perhaps 
connected with the presence of a thin film of adsorbed 
oxygen, which effectively seals off the surface of the 
metal from reaction. 

The reactive form of molybdenum necessary for 
the reaction with carbon monoxide was produced in 
the reaction vessel by the reduction of molybdenum 
pentachloride by a suitable reagent, such as metallic 
zinc, in the presence of a water-free solvent, such as 
ether. The two-step process for the formation of car¬ 
bonyl may be represented by equations (1) and (2). 

2MoC 1 5 + 5Zn—> 2Mo (active) + 5ZnCl 2 (1) 
Mo (active) + 6CO —> Mo(CO)6. (2) 

The mechanism of this process and the exact course 
of the formation of the carbonyl have been the subject 
of considerable discussion in the chemical literature, 
but the above reactions seem to fit the observed facts 
in as simple a way as possible. 

21.2.2 Preparation of 
Anhydrous Molybdenum Chloride 

Molybdenum chloride (M0CI5), as was stated 
above, is one of the principal reagents in the prepara¬ 
tion of the carbonyl. Because this reagent was not 
available in quantity, part of the investigation was 
perforce the working out of a method for the manu¬ 
facture of this compound in pound lots or larger 
quantities. 

In the past, anhydrous molybdenum pentachlor¬ 
ide, the only stable known chloride of this element, 
has been prepared in gram quantities by the reaction 
of a slow current of dry chlorine with finely powdered 
molybdenum metal, molybdenum oxide (in this in¬ 
stance carbon tetrachloride vapor has sometimes 
been substituted in part for chlorine), or molyb¬ 
denum sulfide. The chlorine, which is diluted with an 
inert gas, is passed over the powdered metal or com- 


CON FIDENTIAL 




PREPARATION OF MOLYBDENUM CARBONYL 


421 


pound in a hard glass or quartz tube at elevated tem¬ 
peratures. The molybdenum pentachloride has then 
been removed from the reaction zone, partly by sub¬ 
limation and partly by the gas stream, and preserved 
in sealed glass tubes, as it is readily oxidized by moist 
air. 

A new method developed for the preparation of the 
pentachloride involves directing a jet of chlorine 
taken from a tank of the commercial gas upon the 
surface of powdered molybdenum metal. The reac¬ 
tion takes place readily without external heating, and 
the molybdenum pentachloride melts and can be 
drawn off through a valve as needed. In practice, the 
reaction vessel is equipped with an external heating 
device to sublime the pentachloride from any residual 
metal or impurities, and it is condensed in a second 
vessel. Ordinary commercial molybdenum metal pow¬ 
der gives pentachloride of ample purity. The yield is 
quantitative, and the properties of the product agree 
with those recorded in the literature. 

An apparatus that produced molybdenum penta¬ 
chloride at the rate of over 2 lb per hour is shown in 
Figure 1. On the basis of the experience gained in 


constructing and operating jthis apparatus, it should 
prove a simple engineering job to build a continu¬ 
ously operating commercial unit capable of produc¬ 
ing molybdenum pentachloride of adequate purity 
for any purpose in amounts ranging from 10 to 100 
lb per hour. However, on the basis of the work so far 
carried on, no estimate of the cost of molybdenum 
pentachloride when prepared in large quantities is 
possible. 

212 3 Procedure for 

Producing Molybdenum Carbonyl 

Preparation of Crude Product 

Molybdenum carbonyl was formed by the reaction 
of carbon monoxide with the molybdenum resulting 
from the reduction of anhydrous molybdenum chloride 
in the presence of a water-free solvent. The prepara¬ 
tion of the carbonyl was carried out at high pressure 
and with constant and effective agitation to provide 
prolonged contact with the gas and also to provide 
fresh surfaces for the reaction. After many changes in 



tir" 

Figure 1 . Chloridizing apparatus used for improved method of preparing molybdenum pentachloride. Right to left: 
chlorine tank, safety flasks, asbestos-covered flask containing molybdenum, flasks for trapping fumes, container for 
pentachloride. (Figure 5 in NDRC Report A-422.) 


CONFIDENTIAL 










422 


PLATING OF MOLYBDENUM, TUNGSTEN, AND CHROMIUM 


apparatus and operating conditions had been made, 
the production of purified molybdenum hexacarbonyl 
in lots of 1 lb or more, with a yield of over 85% of the 
theoretical, was successfully carried out. The rate of 
preparation was increased to over 2 lb per hour. 

A forged steel autoclave, copper plated on the in¬ 
side, of 1-gal capacity, shaken in a horizontal plane 
with a 6-in. stroke by a motor running at 180 rpm 
was used as a reaction vessel. The autoclave and the 
horizontal shaker are illustrated in Figure 2. 



Figure 2. One-gallon autoclave in horizontal shaker. 
(Figure 8 in NDRC Report A-422.) 


A discussion of the reagents and the conditions 
used in the preparation of the carbonyl follows. 

The molybdenum pentachloride, the preparation 
of which was described in Section 21.2.2, was used in 
the form of irregular, broken lumps, about the size of 
lump sugar, of material which had been fused. Pre¬ 
liminary brief exposure of this material to the air 
seems to assist the reaction, probably by allowing ex¬ 
cess chlorine to evaporate. 

Anhydrous ether was found to be the best solvent. 
The zinc, in slight excess over the amount theoreti¬ 
cally required, was added to the ether solution of the 
molybdenum pentachloride in the form of a powder 
fine enough to pass a 30-mesh screen. It was found 
advisable to add the zinc in one mass, wrapped in 
paper, so that the reaction would start slowly. If the 
zinc is added in larger pieces or is placed bare in the 
pressure vessel, the carbonyl first formed tends to 
coat its surface, and the reaction nearly stops. Zinc 
powder of a coarser grain requires a much longer time 
to complete the reaction than does the 30-mesh pow¬ 
der. Dispersing the zinc on glass wool before introduc¬ 
ing it into the container reduces the tendency of the 
system to seal off the surface of the zinc, but makes 


it more difficult to distill the product. Aluminum 
powder, magnesium grains, copper filings, and hy¬ 
drogen gas were also tried as possible reducing agents, 
but were less effective than zinc. 

A pressure of carbon monoxide of about 700 psi 
seems to be the most effective. At pressures of 600 to 
650 psi the rate of the reaction and hence the yield 
in a given time are considerably reduced. An increase 
in pressure to 800 psi seems to have no beneficial 
effect, but this point needs further study. 

A special jacketed 5-gal capacity autoclave, cop¬ 
per plated on the inner surface, equipped with an 
automatic stirring device capable of running at high 
and variable speeds, and having charging and dis¬ 
charging vents at top and bottom, was built for further 
experimental production on a larger scale. It was 
thought advisable in this case to use a jacketed vessel 
so that external heating or cooling could be applied 
to increase or moderate the rates of the reactions and 
possibly to increase the yield while cutting down the 
time. In the production of small quantities of the car¬ 
bonyl, the reactions generate enough heat to maintain 
a reasonable rate, but not enough to cut down the 
yield of product by the occurrence of side reactions. 

Due to the termination of the program the 5-gal 
autoclave was not used. On the basis of the produc¬ 
tion of about lj^ lb of purified carbonyl per success¬ 
ful run of the 1-gal vessel, a production of 6 to 8 lb 
per run of 5 to 6 hr would be a reasonable expectation 
on this larger scale. 

Removal of Ether 

When the reaction was reasonably complete, as 
shown by the constancy of pressure of carbon monox¬ 
ide on the gauge over a period of some minutes, the 
gas was shut off, the pressure vessel opened, and the 
contents removed. It was found that two phases were 
present: a liquid phase, consisting of an ether solu¬ 
tion of molybdenum carbonyl, zinc chloride, and un¬ 
reacted molybdenum chloride; and a solid, or more 
strictly a semisolid phase, which consisted of a mix¬ 
ture of all the above with small lumps of reduced 
molybdenum metal which had not reacted with the 
carbon monoxide. In later preparations, where the 
reaction had gone more nearly to completion, the 
separation into two phases took place much more 
slowly, and it was found simpler and quicker to distill 
the ether from the entire mass without separation. 

The contents of the reaction vessel were transferred 
to a glass flask of suitable size, and the ether distilled 


CONFIDENTIAL 










PREPARATION OF MOLYBDENUM CARBONYL 


423 


off. Water was added drop wise from a funnel with a 
glass stopcock in the stem, to keep zinc chloride in 
solution and to permit the removal of the last traces 
of ether without overheating. If the water is added 
rapidly, the temperature rises quickly and molyb¬ 
denum carbonyl is lost. 

To prevent loss of carbonyl at the present stage, 
and also to dry the ether, a reflux condenser, consist¬ 
ing of a wide tube filled with lumps of anhydrous cal¬ 
cium chloride, was attached to the top of the still, and 
the virtually dry, carbonyl-free ether vapor passing 
from this was condensed by a water-cooled conden¬ 
ser. This recovered ether, after redistillation through 
an ordinary reflux condenser, was found to be as 
effective a solvent for molybdenum pentachloride as 
was fresh ether. About 85 to 90% of the ether can 
thus be recovered. 

Steam Distillation 

The molybdenum carbonyl was separated from the 
other constituents by steam distillation. In this man¬ 
ner it could be recovered pure, but not dry, without 
decomposition by overheating. The distillation rate 
was necessarily slow, about 250 g/hr, since molybde¬ 
num carbonyl begins to decompose at about 150 C. The 
apparatus, shown in Figure 3, was simple. Water was 
boiled in the flask at the right, the residue from the 
ether distillation was continuously introduced by a 
screw-feed device into the center flask, which was 
equipped with a powerful stirrer as the mass was 
sticky, and the carbonyl was condensed in the flask 
at the left, which was cooled in cracked ice. The flask 
at the extreme left was used to condense any car¬ 
bonyl vapor’that escaped the main condensing flask. 

Drying and Sublimation 

The molybdenum carbonyl, which carried 35 to 
40% water after the steam distillation, had to be care¬ 
fully and completely dried, before it could be used in 
the plating experiments described in Section 21.3. The 
effect of water on the stability of the carbonyl is not 
great, but the wet mass tended to become sticky and 
difficult to handle. After considerable experimenta¬ 
tion, the following process, which gave a completely 
dry product with virtually no loss, was devised, using 
the apparatus shown in Figure 4. 

The flask containing the moist carbonyl was in¬ 
verted to drain off most of the water. It was then 
placed in an electrically heated oven and the closed 


end of a wide glass tube of suitable size was passed 
through the neck of the fla^k, the joint being made 
tight by a rubber stopper, through which also passed 
a smaller tube connected to an effective vacuum 
pump. The pump was then started, without heating 
the flask containing the carbonyl, and the remainder 
of the water was evaporated by the difference in 
pressure. The water was condensed in a chilled re¬ 
ceiver placed between the carbonyl flask and the 
pump. 

After the water had been thus pumped off, a dry 
receiver was inserted in the line and the carbonyl 
sublimed onto the walls of the large tube inserted in 
the flask. This was accomplished by heating the oven 
to 100-105 C while the pressure was kept at a few 
millimeters of mercury and the inside of the large 



Figure 3. Apparatus for steam distillation of molyb¬ 
denum carbonyl. (Figure 12 in NDRC Report A-422.) 


tube was cooled by a stream of water. The receiver 
outside the oven was cooled in cracked ice and served 
as a trap to collect any carbonyl that passed the main 
condenser. Nearly 700 g of sublimed carbonyl was 
collected from one lot by this process. 

The sublimed molybdenum carbonyl is preserved 
in glass-stoppered bottles sealed with paraffin wax to 
exclude all moisture. The product is apparently not 
sensitive to light, but it is probably well to keep it 
stored away from direct sunlight. 

Properties of Sublimed Product 

The properties of molybdenum hexacarbonyl thus 
prepared may be summarized as follows: 

Formula: Mo(CO)6 
Molecular weight: 264 
White crystalline solid 
Specific gravity at 25 C: 1.96 


CONFIDENTIAL 







424 


PLATING of molybdenum, tungsten, and chromium 


On heating in air at 1 atm pressure, decomposes at 
about 150 C without melting. Sublimes at lower 
temperatures with little decomposition. 

Vapor pressure at 25 C: about 0.1 mm Hg. 

Soluble in ethyl ether to the extent of about 2% by 
weight. 

Insoluble in water and mineral acids. 

Stable in the presence of water and acids up to 
about 150 C. 

Decomposes in strongly basic solutions in the pres¬ 
ence of the halogens. 

Little is known as to the toxicity of molybdenum 
carbonyl. The more highly volatile, liquid or gaseous 
carbonyls of iron and nickel are extremely active 
poisons. Molybdenum hexacarbonyl vapor is reported 
to be relatively nonpoisonous in low concentrations, 
but this statement needs confirmation. Work on this 
point is reported to be in progress, but the results are 
not known to us. The solid appears to be noncorrosive 
to the skin, but in the present state of our knowledge, 
it should not be handled without gloves. There is 
always the possible risk of dermatitis if handled over 
an extended period. The substance is undoubtedly 
poisonous if swallowed but the dangerous and lethal 
doses are not known. It has been reported that an 
ether solution of molybdenum carbonyl may deto¬ 
nate on standing. There are many precautions to be 
taken with respect to the various reagents and prod¬ 
ucts involved in the preparation of molybdenum 
carbonyl. 


Summary and Recommendations 

A generalized flow sheet for the production of 
molybdenum carbonyl by the process just described 
is given is Figure 5. 

Much work remains to be done on the most effec¬ 
tive means for introducing the reagents into the proc¬ 
ess and on the most suitable and effective apparatus 
for carrying out the reactions involved and for purify¬ 
ing the product. Some further experiments on the 
most effective conditions of temperature and pressure 
for carrying out the successive steps are required. 

Apparatus has been designed and built for the 
production of molybdenum carbonyl in lots of 5 to 
10 lb. Should the development of further applications 
for the plating of molybdenum by thermal decompo¬ 
sition of the vapor of its carbonyl warrant it, engi¬ 
neering studies leading to the preparation and puri¬ 
fication of the carbonyl on a suitable scale should 
present no undue difficulties, although they may 
need to be extensive. 

2 * 3 VAPOR-PHASE PLATING OF 
MOLYBDENUM 93 

21,31 General Procedure 

Vapor-phase plating of molybdenum is the pro¬ 
cedure whereby molybdenum is deposited on a hot 
surface by the thermal decomposition of molybdenum 



Figure 4. Apparatus for drying and sublimation of molybdenum carbonyl. (Figure 13 in NDRC Report A-422.) 

CONFIDENTIAL 









VAPOR-PHASE PLATING OF MOLYBDENUM 


425 



Figure 5. Flow-sheet for the production of molyb¬ 
denum carbonyl. (Figure 14 in NDRC Report A-422.) 


carbonyl. An apparatus was developed whereby very 
pure hydrogen at a low pressure passing over crystal¬ 
line molybdenum carbonyl held at a constant tem¬ 
perature slightly above room temperature carries 
molybdenum carbonyl vapor into an electric furnace 
containing the object to be plated which has been 
uniformly heated to a predetermined temperature. 
The carbon monoxide formed by the reaction is re¬ 
moved by pumping at the end of the plating chamber. 

Plating may be carried out in the absence of hydro¬ 
gen ; however, there are important advantages in the 
use of this gas before plating and as a constituent of 
the plating gas. Heating in dry hydrogen before plat¬ 
ing can be effective in removing oxide films from the 
surfaces to be plated. As a constituent of the plating 
gas, hydrogen increases the flow of carbonyl with 


increase in partial pressure of the latter. 

The addition of water to the plating gas is benefi¬ 
cial in some cases, as is described later. In order to 
add water to the hydrogen of the plating gas, only a 
very simple modification in the apparatus is neces¬ 
sary. A by-pass makes it possible to admit dry hydro¬ 
gen instead of wet hydrogen to the carbonyl chamber. 

A schematic diagram of the apparatus is shown in 
Figure 6. The small circles on this diagram represent 
vacuum-tight valves; those which are marked with 
the letter N are needle valves used for adjustment of 
the hydrogen flow. In all cases the hydrogen must be 
very carefully purified, preferably by diffusion through 
palladium. The carbonyl vapor is released from its 
chamber at a constant rate, which is determined by 
the temperature of the chamber and by the rate of 
flow of the dry or wet hydrogen. 

The plating chamber contains means for support¬ 
ing the specimen, for directing the plating gas to 
selected areas of the specimen if necessary, and for 
heating the specimen. The design of the chamber 
depends on the size and shape of the specimen to be 
plated. When a gun liner (Section 21.3.4) was being: 
plated, it was held fixed and a water-cooled carbonyl 
injector moved uniformly within it so that no increase 
in concentration of the reaction products took place 
at any point. The plating chamber in this case was a 
quartz tube and the liner was heated by high-fre¬ 
quency induction. Injector and liner were mounted 
vertically. Chromel-alumel thermocouples attached 
to the liner enabled the operator to know its tempera¬ 
ture, which had to be carefully controlled. 

The pump, which is used to remove the products 
of the reaction, maintains pressure gradients in the 
carbonyl and plating chambers and also helps to 
regulate the pressure in the plating chamber. A 
McLeod gauge is used to read the pressure behind the 
injector or at the plating chamber. The reaction 
products, which are removed from the chamber by 
pumping, are collected in a liquid air trap, which is 
designated LA in Figure 6. 

21,3,2 Properties of Molybdenum Plates 

Introduction 

The properties of a molybdenum plate, of course, 
depend on the plating conditions. The effects of vary¬ 
ing the procedure were studied by means of test 
specimens. Only after the plates were found to have 
desirable properties was the procedure used for prac- 


CONFIDENTIAL 

















































426 


PLATING OF MOLYBDENUM, TUNGSTEN, AND CHROMIUM 



Figure 6. Schematic diagram of apparatus for vapor-phase plating of molybdenum. (Figure 5 in NDRC Report A-421.) 


tical applications. Mcfet of the experiments discussed 
here, therefore, are those that were carried out on 
test specimens. 

In a consideration of the properties of a coating or 
plate deposited on a metal by the thermal decomposi¬ 
tion of a carbonyl, distinction must be made between 
the properties of the coating itself and the nature of 
the metal-to-plate bond, as with any other coating. 
In this section only the former are considered. They 
include: 

Chemical composition and crystal structure of the 
plate; 

Mean crystal size and preferred orientation, if any; 

Hardness; 

Mechanical strength; 

Texture, if plate is nonisotropic. 

Chemical Composition of the Plate 

The dominant reaction products formed by the 
decomposition of molybdenum carbonyl are molyb¬ 
denum and carbon monoxide. As was stated in Sec¬ 
tion 21.1.2, there are secondary reactions which may 
take place depending on the plating conditions. The 
carbon monoxide may decompose to yield carbon and 
carbon dioxide. The carbon may react with the mo¬ 


lybdenum to form a carbide or it may be deposited 
as interstitial carbon within the coating. The kind of 
plate produced thus depends on the factors which 
determine the course of the secondary reactions when 
the equilibrium is shifted. 

Plating experiments in which carbonyl vapor was 
the only gas admitted to the plating chamber showed 
how the composition of the plate that was deposited 
on an iron disk varied with the temperature of the 
specimen and the pressure of the carbon monoxide 
from the decomposition of the molybdenum carbonyl. 
The results are summarized in Table 1. The nature 
of the plates was determined from x-ray diffraction 
patterns obtained from the plated surface. Both car- 


Table 1. Nature of plates deposited from molybdenum 
carbonyl vapor. 


Plating 

temperature (C) 

Low carbon 
monoxide pressure* 

High carbon 
monoxide pressure* 

200-300 

Cubic carbide 

Cubic carbide 

300-400 

Cubic carbide 

Cubic carbide 

400-500 

Molybdenum 

Cubic carbide 

500-800 

Molybdenum 

Hexagonal carbide 


* The pressure range from 0 to about 0.2 mm is described as “low carbon 
monoxide pressure,” and the range from about 0.2 to about 10 mm as 
“high carbon monoxide pressure.” 


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VAPOR-PHASE PLATING OF MOLYBDENUM 


427 


bides have the formula M02C. It was necessary to 
determine the empirical formula in the case of the 
cubic carbide by chemical analyses, for this carbide 
of molybdenum had not previously been reported. 
Results of other experiments show that in all cases 
where any plate at all is deposited, the carbon mon¬ 
oxide enters into the reaction to some extent; that is, 
all plates carry either some carbide of molybdenum 
or some interstitial carbon deposited as such in the 
plate. 

The addition of hydrogen to the plating gas greatly 
assists in reducing oxide coatings on the metal, with 
a resultant tendency to strengthen the bond. The 
addition of hydrogen does not tend to lower the car¬ 
bon content of the plate unless the ratio of hydrogen 
to carbon monoxide is very large, as is shown in Table 
2. The geometry and the temperature of the plating 
chamber have more effect in this respect. 


Table 2. Plating conditions at the boundary between 
molybdenum metal and cubic molybdenum carbide, 
hydrogen present. 


Plating 

Partial 

pressure 

Plate 

temperature (C) 

II 2 (mm) 

CO (mm) 

formed 

425 ' 

0.128 

0.080 

Cubic carbide 

400 

0.180 

0.054 

Cubic carbide 

325 

0.950 

0.026 

Molybdenum 


At high temperatures and at high pressures of 
residual carbon monoxide and of hydrogen, the mix¬ 
ture of molybdenum and molybdenum carbide forms 
an unattached, very fine, powdery layer in the cham¬ 
ber, due to decomposition of molybdenum carbonyl 
in the gas phase. This tendency sets an upper limit to 
plating rates. By a suitable adjustment of conditions 
almost any mixture from 100% metallic molybdenum 
to 100% hexagonal molybdenum carbide can be 
produced. 

If the hydrogen is mixed with water vapor before 
being passed into the system, lower carbon contents 
of the plates can be obtained, and the plating rate 
greatly increased. This effect cannot be obtained by 
admitting moist air into the system, as oxidation 
takes place at once, and the plate consists largely of 
unadherent molybdenum oxide. 

Various modifications of the “normal” plating pro¬ 
cedure, i.e., molybdenum carbonyl vaporized in a 
stream of dry or wet hydrogen, were tried. The effect 
of adding carbon dioxide to the plating gas to de¬ 
crease the formation of carbon by shifting the equi¬ 
librium 2CO ^ C + CO2, seems to be very small 


at 500 C, above which temperature the quality' of the 
plate begins to fall off. 

Hydrogen sulfide decreases the formation of carbon 
more effectively than does hydrogen, at low tempera¬ 
tures, but the adherence of the plate and the quality 
of the bond suffer greatly. All traces of hydrogen 
sulfide must be swept most carefully out of the appa¬ 
ratus before a subsequent hydrogen reduction of the 
metal surface. At low concentrations of hydrogen 
sulfide (pressures of 0.006-0.0002 mm) the plates 
formed are of molybdenum metal. When this gas is 
added in large excess, plates of molybdenum sulfide, 
MoS 2 , are formed. These appear to be adherent, but 
are, of course, very soft. Molybdenum sulfide in 
nature is about as soft as natural graphite. 

The addition of air plus excess oxygen to pure 
carbonyl in the correct proportions results in a thin, 
brittle, adherent plate of molybdenum oxide, M0O2. 
Thick deposits of this type cannot be produced, for 
large amounts of air produce a nonadherent coating. 

Physical Characteristics of the Plates 

Crystallographic Orientation. Where any degree of 
adherence is obtained between plate and metal, the 
plate consists of an interlocking mass of oriented 
molybdenum crystals. The exact crystallographic 
orientation seems to vary with the chemical char¬ 
acteristics of the plate and with the plating condi¬ 
tions. In general the softer the plate, the larger and 
more brilliant are the crystals. Figure 7 shows the 
large, brilliant crystals typical of a soft plate. The 
preferred direction of crystal growth appears to be, 
in nearly all cases, normal to the surface of deposition. 
This is particularly true when plates of considerable 
thickness are formed. The tendency towards columnar 
structure, illustrated in Figure 8, is less in plates with 
very small individual crystals, which can apparently 
only be obtained where there is either considerable 
interstitial carbon but no carbide, or in those plates 
containing no molybdenum metal but only carbide 
plus excess carbon. 

Hardness. The hardness of the molybdenum plates 
increases with the carbon content, whether it is pres¬ 
ent as carbide or as interstitial carbon. Brittleness of 
the harder plates, on the other hand, increases with 
an increase in the proportion of carbide. When a 
hard, but not brittle, plate is desired, therefore, plat¬ 
ing conditions should be such that carbon is deposited 
interstitially and not as carbide. When no carbide is 
present, the hardness increases with decreasing crystal 


CONFIDENTIAL 








428 


PLATING OF MOLYBDENUM, TUNGSTEN, AND CHROMIUM 


Table 3. Hardness of certain molybdenum plates. 


Plating 

temDer- 

Percentage 

composition 


ature 

(C) 

Vickers 

hardness 

Cubic 

carbide 

Metal 

Notes 

380 

1460 

100 

0 

Large oriented 
crystals 

410 

1300 

85 

15 

Large oriented 
crystals 

410 

1550 

15 

85 


450 

1560 

0 

100 

Very small 
crystals 

540 

940 

0 

100 

Small crystals 

475 

1550 



As plated 

475 

880 

same 



After 1 hr at 775 
C in wet H 2 

475 

plate 8g0 



After 1 hr more 
at 750 C in wet 

h 2 


size, as shown in Table 3, as well as with increasing- 
carbon content. There seems to be no doubt that the 
presence of interstitial carbon is the cause of both the 
small crystal size and the hardness. 

Below 500 C the use of wet hydrogen has little ef¬ 
fect on the hardness; at higher temperatures an in¬ 
crease in water content of the plating gas lowers the 
hardness considerably because the carbon content of 
the plate is lowered. The presence of small amounts 
of water in the plating gas above 500 C shifts the re¬ 
action for the formation of carbon from carbon mon¬ 
oxide in the direction of the latter. Removal of pre¬ 



Figure 7. Photomicrograph of molybdenum plate 
consisting of highly oriented crystals showing well de¬ 
veloped (100) faces. The (111) planes are parallel to the 
general surface. ( 875X .) (Figure 8 in NDRC Report 
A-421.) 


cipitated carbon takes place in accordance with the 
shift in equilibrium when plates are softened by an¬ 
nealing at high temperatures in wet hydrogen as 
shown in Table 3. 

Annealing at high temperatures in dry hydrogen 
also softens molybdenum plates. A plate with high 
carbide content, however, is not as effectively soft¬ 
ened as one with a low carbide content. It is possible 
to prepare plates as soft as 200 Vickers by plating in 
wet hydrogen at 600 C and then giving the plate a 
high-temperature anneal. 



Figure 8. Photomicrograph of cross section of a 5-mil 
molybdenum plate showing large crystals and well- 
developed columnar structure. ( 500X .) (Figure 9 in 
NDRC Report A-421.) 

213 3 Plate-to-Metal Bond 

Introduction 

The strength of the bond between metal and plate 
is especially important in the case of these pyrolytic 
plates, sfiice the coefficients of expansion of the metals 
concerned differ widely (steel about 13 X 10 -6 ; mo¬ 
lybdenum, 5.5 X 1(R 6 ; tungsten, 4.5 X 10 -6 ). Vari¬ 
ous expedients have been tried to obtain a firm bond, 
without complete success, but for devices where 
severe mechanical shock and thermal shock are ab¬ 
sent, or are not present simultaneously, the results of 
experiments to date appear promising. 

In general, a good bond requires a metal surface 
that is perfectly clean, especially free from oxide. The 
surface must be kept clean until it is adequately pro¬ 
tected by plate. For small surfaces, cleaning with 
very pure abrasive under absolute alcohol can be 


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VAPOR-PHASE PLATING OF MOLYBDENUM 


429 


used, but the best general method appears to be the 
exposure of the surface to a stream of hydrogen gas 
at high temperatures in the plating chamber before 
passing in the plating gas. 

The temperature and length of exposure needed 
depend primarily on the ease of reducibility of the 
oxide coating. Thus alloy steels and stellite require 
longer cleaning times and higher temperatures than 
do pure iron or copper. In some cases it becomes 
necessary to heat the steel surface to temperatures 
which are above the transition point for periods that 
are long enough to change the structure and proper¬ 
ties of an underlying layer of measurable thickness, 
thus making it necessary to heat-treat the article 
after plating. Stellite surfaces need to be heated to 
1000-1100 C for Yi hr for complete deoxidation. 

Formation of Good Bonds 

The best procedure to give an adequate bond in¬ 
volves a preliminary cleaning operation, followed by 
hydrogen reduction, after which the plating is started 
at a lower (but still fairly high) temperature and at a 
low rate. Conditions to be chosen vary with the chem¬ 
ical nature of the articles to be plated and their 
geometry. 

Testing Bonds 

In order to save the time and expense required for 
full “use” tests of the strength of the plate-to-metal 
bond, various laboratory tests were developed. A 
very poor bond fails by spalling when cold, that is, by 
separating from the underlying metal in flakes as the 
result of the strains set up on cooling to room tem¬ 
perature. Bonds that withstand this simple test some¬ 
times fail when the article is slightly bent or gently 
hammered. Failure usually takes place first at a sharp 
edge. 

A test was developed especially for use on speci¬ 
mens prepared for gun barrel liners. In this test a ring 
Yi in. in inside diameter and in. long, with lands 
and grooves on the inner surface, was crushed in a 
compression-type testing machine. It was possible to 
plate such rings so that the bond would withstand 20 
per cent to 30 per cent of axial compression without 
failure. As the investigation progressed, plates that 
could withstand 20 per cent axial compression would 
be regularly prepared, and only articles prepared by 
techniques that passed this test were subjected to use 
tests. 


A quantitative shear test, using a precision shear¬ 
ing tool actuated by a measured weight attached to a 
lever arm beyond the tool, and so arranged that the 
tool always assumed the same angle of shear, was also 
developed. This device indicated the average value of 
the strength of the plate perpendicular to the bond 
(best plates as shown by crushing test) to be 100,000 
psi; of the bond, 54,000 psi; and of the hardened and 
heat-treated steel ring, 87,000 psi. 

Effect of Heat Treatment 

In the course of these experiments it was shown 
that heat treating subsequent to plating injures the 
bond, as evidenced by the severe spalling shown in 
Figure 9. The damage to the bond was apparently in 
direct ratio to the length of time of the subsequent 
heating. 

Both heat treating and crushing appear to convert 
interstitial carbon to the hexagonal carbide, Mo 2 C. 
Furthermore these treatments appear to form a diffu¬ 
sion layer of the carbide at the bond interface. As 
molybdenum carbide is known to be brittle, this 
would account for the failure. This diffusion layer is 
also formed when the use to which an article is put 
subjects it to similar conditions. It has never shown 
up in photomicrographs taken of specimens in the 
“as-plated” condition. Experiments show that 800 C 
is the maximum temperature that a plate-to-steel 
bond can safely stand, and about 950 C is the safe 
maximum for a plate-to-stellite bond. 

Other Studies of the Bond 

Attempts were made to improve the bond by alter¬ 
ing the surface layer of a steel article, and also by 
applying electroplates of various metals to the steel 
or stellite surface before the pyrolytic plating. De- 
carburizing a steel surface to a depth of 30 mils, 
either with or without subsequent nitriding, had no 
effect. Electroplating steel specimens with copper, 
gold, gold on cobalt, gold on nickel, chromium on 
cobalt, platinum bright on nickel, and palladium on 
cobalt, did not improve the resistance of the bond to 
heat treatment at 900 C. 

The greatest improvement was shown by molyb¬ 
denum coatings that were plated on steel specimens 
which had previously been plated with nickel or with 
platinum on nickel. However, the improvement 
shown on flat specimens was not repeated when rifled 
tube sections were electroplated thus. It may be noted 


CONFIDENTIAL 



430 


PLATING OF MOLYBDENUM, TUNGSTEN, AND CHROMIUM 



Figure 9. Crushed halves of a molybdenum-plated steel ring. The upper half was crushed 37 per cent without spalling 
or cracking of the plate. The lower half was heated for 30 minutes at 900 C after plating and was then crushed 12 per cent, 
which resulted in very severe spalling of the plate. (Magnified 4X.) (Figure 12 in NDRC Report A-421.) 


that it is difficult to obtain satisfactory electroplates 
on such a shape. No other reason for this difficulty is 
apparent. On stellite surfaces, an undercoat of elec¬ 
troplated nickel gave improvement in resistance to a 
heat treatment at 900 C, but failed after heating for 
1 min at 1100 C. 

213 4 Application to Gun Liners 

By far the greatest amount of time spent on the 
experimental plating program, after the fundamentals 
had been worked out, was devoted to the problem of 
molybdenum-plating short liners of steel and stellite 
for tests in caliber .50 gun barrels. Many liners were 
plated and 42 were fired, some in a caliber .50 aircraft 
barrel and others in the caliber .50 erosion-testing 
gun. The use of these guns for tests of gun liners is 
described in Sections 11.2.2 and 11.2.1, respectively. 

Plating a bore of such small diameter is a more 
difficult job than in the case of a larger one, partly 
due to the greater impedance of the gas flow by walls 
and partly due to the difficulties of building a “plat¬ 
ing-head” or injector that will function in a small 
tube. For the liners used, the maximum plating rate 
at which good adherence could be obtained was about 
0.18 mil/hr. Calculations indicate that increasing the 
pumping speed would not increase this rate appre¬ 


ciably, as it appears to be about 75 per cent of the 
rate possible with a pump of infinite capacity. 

Electrolytic polishing of the liner bore under care¬ 
fully controlled conditions gave a satisfactory surface 
for plating. This etching process was followed by 
preliminary reduction of the surface in hydrogen; 
preliminary heat treatment in hydrogen, to be sure 
that pure hydrogen was the last thing in contact with 
the metal before plating; and the application of the 
first, or bond, layer of plate at a lower rate than for 
the balance of the plating. 

The best performance given by a molybdenum- 
plated stellite finer fired in the. erosion-testing gun 
was that of a rifled finer with an 8-mil thick, soft-type 
plate.™ Wet hydrogen was added to the plating gas 
to yield this type of plate. This liner was fired 240 
rounds at a velocity of about 3,G50 fps, a pressure of 
58,200 psi, and a rate of 4 to 6 rounds per minute. 
Ball bullets were used with double-base powder con¬ 
taining 20% nitroglycerin. With I MR powder under 
approximately the same hypervelocity conditions, 
the same maximum performance was obtained. The 
best performance for a steel liner under the hyper¬ 
velocity conditions noted above was 230 rounds with 
IMR powder. 

In all cases failure was due to spalling of the plate. 
By far the greatest amount of spalling occurred at 


CONFIDENTIAL 





VAPOR-PHASE PLATING OF MOLYBDENUM 


431 


the land edges, where the mechanical forces were 
greatest. Areas where the plate remained intact 
showed virtually no scoring or powder gas erosion. 
The facts that failure occurred under hypervelocity 
conditions at about the same number of rounds for 
both double-base and single-base powder and that 
spalling was greatest on land edges indicate that 
failure was due to mechanical rather than to chemical 
conditions. From knowledge of the chemical behavior 
of molybdenum and from the experience with mas¬ 
sive molybdenum liners, discussed in Section 18.6, 
this was to be expected. 

There seem to be two dominant causes for the 
failure of the plates. d First, the maximum possible 
thickness for adherent plates is still not thick enough 
to keep the bond from being weakened by the heat 
developed, the bond becoming brittle as the result of 
the formation of a diffusion layer during firing. Sec¬ 
ond, the plate tends to split between the parallel- 
oriented, columnar crystals and thus leaves the basic 
metal open to attack by the powder gas. The cracks 
tend to pass down into a steel liner, and pieces of steel 
“‘chunk” out. When the hot gases from double-base 
powder reach a stellite liner, melting starts at once so 
that flakes of plate are undermined and detached 
from the surface. Also, the hot powder gases can react 
with the exposed surface of a steel liner to yield 
brittle products, described in Chapter 12, which are 
removed, together with molybdenum plate, by the 
mechanical action of the projectile. When once the 
plate fails at a few spots, the action becomes cumu¬ 
lative. The photomicrographic evidence for the cause 
of the failure of the plate-to-steel bond seems a little 
clearer than does the evidence of failure for the plate- 
to-stellite bond. 

The effort to make a satisfactory plated gun liner 
by using a thicker plate had two other difficulties. 
One was that the plating time became inordinately 
large. The other was that the rifling lost its definition 
when the plate was thicker than 10 mils in a caliber 
.50 liner having rifling 10 mils deep. Some attention 
was given to the possibility of plating a smooth-bore 
liner and rifling it after plating. This procedure would 
have required a softer plate that could subsequently 
be hardened. Reconnaissance experiments to produce 
such a plate were unsuccessful. 

In summary, then, it may be pointed out that the 
requirements for a molybdenum plate to be applied 


d Examinations of the fired liners were carried out at several 
different places. 76 * 77 ' 93 


to gun liners by the vapop-phase process are as fol¬ 
lows : First, a type of plate that is soft enough when 
deposited so it can be rifled by machining, and then 
can be hardened by heat treatment or in some other 
way; second, a type of plate that will have a non- 
oriented crystallographic structure; and third, a plat¬ 
ing rate of at least one mil per hour, so that a plate at 
least 15 mils thick can be deposited in a reasonable 
time. 

213 5 Other Applications 

Rocket Nozzles 

Rocket nozzles are subject to powder-gas erosion 
but not to mechanical shock. Since molybdenum has 
superior resistance to attack by powder gases and, as 
liners, fails in guns only because of mechanical shock, 
it was thought worth while to plate rocket nozzles 
with molybdenum. Several copper and steel nozzles 
of various designs were plated by the pyrolytic 
method. 45 

A special plating apparatus had to be designed and 
built, since the nozzles had tapering orifices and hence 
the “plater’ 7 designed for cylindrical gun liners would 
not give plates of the desired thicknesses. Plating 
conditions were worked out, and the nozzles thus 
plated stood up, in most cases, to the laboratory 
tests previously described. 

On firing, where the duration of a single blast is up 
to 11 sec and the flame temperature of the powder is 
3000 K, the nozzles failed at the metal-to-plate bond. 
The plate was presumably not thick enough to pro¬ 
tect the underlying copper from melting or the under¬ 
lying steel from being heated above its transition 
point; in addition to this effect, the formation of a 
brittle diffusion layer probably occurred during the 
long-continued heat treatment. The high heat trans¬ 
fer of molybdenum is presumably partly to blame. 
If much thicker plates could be made adherent, a 
greater degree of success might be expected. 

Recoilless Gun Blocks 

Some recoilless gun adjustment blocks were plated 
with molybdenum/ The plates in all cases spalled 
along the leading edge after about 20 rounds. This 

e This work was done at the request of the Allegany Ballistics 
Laboratory of Section H of Division 3, NDRC, and of the 
Explosives Research Laboratory of Division 8, NDRC. 

f This work was done at the request of Frankford Arsenal. 


CONFIDENTIAL 





432 


PLATING OF MOLYBDENUM, TUNGSTEN, AND CHROMIUM 


was not considered to be satisfactory performance. 
Heat transfer, accelerated at the leading edge, was 
the probable cause of failure. A softer and thicker 
plate might give better performance. 

Oxygen Compressor Cylinders 8 

Certain types of oxygen compressor cylinders have 
to work dry, and the wear of the cylinder against the 
carbon piston ring is serious. Under the conditions of 
use-it is the friction between the carbon ring and the 
coating on the inside of the cylinder, rather than the 
wear resistance of the coating, that is being tested. 
Previous tests had shown that hard chromium elec¬ 
troplate gave the best performance. A rather soft 
molybdenum plate lasted 40 hr before wear of the 
carbon ring became serious. The rate of wear against 
the molybdenum plate was only about 8 per cent of 
that against the hard chromium electroplate. No 
other types of molybdenum pyrolytic plates were 
tried, but this appears to be a promising field of ap¬ 
plication of such plates. 

Magnetron Anodes 

Certain electronic devices, such as magnetron 
anodes, dissipate large amounts of heat per unit area, 
the amount increasing with the power radiated. The 
size of such devices is usually fixed; thus if heat dis¬ 
sipation can be increased while electrical conductivity 
can be maintained, an increase in power radiated can 
be obtained. Because of the high rate of heat transfer 
of molybdenum, it was decided to try the effect of a 
soft, pure metal plate on the inside of a copper ring 
which is a portion of the copper anode. It was found 
that a power increase of at least four times over that 
possible with an unplated anode took place. The plate 
remained intact. 

Die Surfaces 

Another application of great promise is to plate the 
surface of dies that have to stand hard wear, but not 
severe shock. A 2-mil molybdenum-molybdenum car¬ 
bide plate, which probably had a hardness above 
1200 Vickers, was applied to a die which consisted of 
a rod at one end of which was a disk for impressing a 
convolute in the grid frame of a vacuum tube, the 

g This application was tried at the request of Division 11, 
NDRC, and the tests of the plates were carried out by a 
Division 11 contractor. 


frame being made of molybdenum and tungsten wires. 
After this die had been used for 50 formings it showed 
no appreciable signs of scoring. The ordinary die shows 
serious deterioration after 30 formings and requires 
refinishing after 100. The test was being continued in 
1946 and it was then estimated that the total useful 
life of the plated die would be many-fold that of the 
ordinary type. 

Another die with a complicated surface, used for 
drawing copper tubing for coaxial cable, was given 
a similar plate. The test of this die was still incom¬ 
plete in 1946, but its performance up to then was 
much better than that of a nonplated steel die. 

Molybdenum Sulfide Plates as Semiconductors 

Molybdenum sulfide plates, which are mentioned 
in Section 21.3.2, appear to present useful possibilities 
as semiconductors. Only preliminary experiments on 
such plates were made. 

213 6 Conclusions and Recommendations 

It appears that the fields of use for pyrolytic plates 
of molybdenum and molybdenum carbide that will 
most probably repay further investigation are those 
where resistance to friction or wear, without mechan¬ 
ical shock, is involved. If considerably thicker plates 
of good adherence can be prepared, there is reason¬ 
able prospect of success in obtaining resistance to 
high temperatures. Such plates would have to be 
thick enough to protect the underlying metal from 
melting and thermal transformation. The columnar 
nature of plates so far prepared increases their lack of 
resistance to the combination of thermal and mechan¬ 
ical shock. Further research may make it practicable 
to prepare plates having an unoriented structure, or 
ones oriented in such a way as to resist the forces that 
have up to now caused failure. 

Inasmuch as a thorough trial had been made of the 
conceivable ways of maintaining the strength of the 
bond during firing, it was concluded in the summer of 
1945 that the pyrolytic plating process as then devel¬ 
oped was not suitable for plating gun liners with 
molybdenum. Furthermore, the liners of massive mo¬ 
lybdenum, described in Chapter 18, had by that time 
lasted over 2,000 rounds in similar tests. The difficulty 
of finding a way to match this result by a liner plated 
with molybdenum by the pyrolytic plating process 
seemed so great as to make further investigation un¬ 
warranted. 


CONFIDENTIAL 




VAPOR-PHASE PLATING OF TUNGSTEN 


433 


Further study of the present type of pyrolytic plate 
should not, however, be dropped. In addition to ex¬ 
tended tests of the more successful applications, 
further research should be conducted on means to in¬ 
crease the plating rate and on methods to increase the 
strength of the plate-to-metal bond, and efforts 
should be made to define more closely the conditions 
for preparing a plate of a given hardness, chemical 
composition, crystal size, or other variable. 


214 VAPOR-PHASE PLATING OF 

TUNGSTEN 93 

21.4.1 Preparation and Properties of 

Tungsten Carbonyl 

For reasons stated in Section 21.1.3, the greater part 
of the efforts on preparing pyrolytic plates for gun 
bores and other purposes was carried out with molyb¬ 
denum carbonyl. Consequently, no detailed investi¬ 
gation was made of processes for the large-scale pro¬ 
duction of tungsten carbonyl or of the raw materials 
from which to make it. Standard laboratory apparatus 
and methods were used to prepare an amount of 
tungsten hexacarbonyl that was sufficient for the ex¬ 
periments that were carried out. 

Anhydrous tungsten chloride is obtainable in com¬ 
merce in quantities adequate for use as a starting 
material for the production of tungsten hexacarbonyl 
on a small scale. If it is needed in larger amounts, it 
can be prepared by the same procedure and with the 
same type of apparatus, with only minor modifica¬ 
tions, as was used for the production of the molyb¬ 
denum compound (Section 21.2.2;. 

Tungsten hexacarbonyl is prepared on a laboratory 
scale in the same fashion as is molybdenum hexacar¬ 
bonyl described in Section 21.2.3. Tungsten hexa- 
chloride is reduced to active tungsten by metallic 
zinc in an ether suspension, and at the same time the 
carbonyl is formed by reaction with carbon monoxide 
under pressure. Upon completion of the reaction, the 
ether is removed by distillation, and the tungsten 
carbonyl is removed from the other substances by 
steam distillation and finally purified by sublimation. 
No studies have been made to determine the most 
favorable conditions for obtaining a high yield, but, 
should it be desired to prepare this compound in 
amounts larger than 100 g at a time, studies such as 
those described in Section 21.2.3 concerning the prep¬ 
aration of molybdenum carbonyl should be made. 


The properties of tungsten hexacarbonyl are virtu¬ 
ally the same as those of the corresponding molyb¬ 
denum compound, except that its vapor pressure at a 
given temperature is about one-fifth that of molyb¬ 
denum carbonyl. Thus at 28.1 C the vapor pressures 
are 0.037 and 0.156 mm, respectively, and at 35.2 C 
they are 0.076 and 0.302 mm. The stability of tung¬ 
sten hexacarbonyl appears to be the same as that of 
molybdenum hexacarbonyl. The toxicity of its vapor 
is presumably the same but the risk of handling it a 
little less because of its slightly lower volatility. 

21,4,2 Properties of Tungsten Plates 

In general, the investigations on the preparation of 
tungsten plates by the pyrolytic, or vapor-phase, 
process stopped at the preliminary stage, since mo¬ 
lybdenum plates seemed to show more promise. Tests 
on improving the bond, on plating tungsten over an 
intermediate plate of another metal, and on the shear 
and crushing strength of tungsten plates were not 
made. 

The same apparatus was used as for the vapor- 
phase plating of molybdenum, ^ described in Section 
21.3.1. When tungsten carbonyl vapor without dry or 
wet hydrogen is introduced into the plating chamber, 
the composition of the plate varies as shown in Table 
4. It seems to be easier to prepare plates of pure tung- 


Table 4. Nature of plates deposited from tungsten car¬ 
bonyl vapor. 


Plating 

temperature (C) 

Low carbon 
monoxide pressure 

High carbon 
monoxide pressure 

200-300 

None 

None 

300-400 

Tungsten 

Cubic carbide 

400-500 

Tungsten 

Tungsten 

500-800 

Tungsten 



sten than of pure molybdenum. The formation of the 
hexagonal tungsten carbide, W 2 C, was never observed 
in these plating experiments. Although no carbon de¬ 
terminations were made, the cubic compound is con¬ 
sidered to be W 2 C by analogy with the cubic phase of 
Mo 2 C that was discussed in Section 21.3.2. 

In general, the tungsten plates are harder than the 
molybdenum plates. Hardness varies from 500 Vick¬ 
ers for plates with very low carbon content, to over 
2000 Vickers for those with high carbon content. 
There is a strong tendency, as in the case of molyb¬ 
denum, for the formation of oriented crystals with the 
long axis normal to the plated surface. 


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434 


PLATING OF MOLYBDENUM, TUNGSTEN, AND CHROMIUM 


Alloy Plates 

Molybdenum-tungsten alloy plates have been pre¬ 
pared by using a finely ground mixture of the two car¬ 
bonyls in approximately' the same mole ratio as the 
ratio of their vapor pressures [Mo(CO)e/W(CO)6 is 
4.5 at 20 C] in the carbonyl chamber. Such plates 
prepared at 600 C in wet hydrogen at a moderate rate 
were quite hard and brittle. Molybdenum plated un¬ 
der these conditions is relatively soft. 

214 3 Tests of Tungsten Plates 

One rifled stellite liner was plated with a 3-mil 
tungsten coating containing some carbide and fired in 
a caliber .50 aircraft machine gun barrel with AP-M2 
ammunition. The plate was removed completely from 
the breech third of this 8-in. liner by firing five 50- 
round bursts at 1-min intervals. This performance 
was not encouraging. 

Tungsten-plated rocket nozzles behaved on test 
the same as did the molybdenum-plated ones, men¬ 
tioned in Section 21.3.5. 

The performance of tungsten-plated adjustment 
blocks in a recoilless gun was no better than that of 
blocks plated with molybdenum (Section 21.3.5). 

Two steel compass bearings were plated as follows. 11 
One received a plate of tungsten high in interstitial 
carbon and the other a plate of tungsten carbide. The 
steel bearings had the shape of a small cylinder with 
a V-shaped recess ground into one end. The hardness 
of the plates was about 2200 microvickers, on which 
scale sapphire is 1800. The ultimate performance was 
reported to be unsatisfactory, probably due to the 
thinness of the plate at the bottom of the bearings 
where the greatest wear would be received. 

No increase in the useful life of a high-speed drill 
was found when a plate of tungsten 1 mil thick was 
applied. 

It should be emphasized that none of the foregoing 
tungsten plates represented the best product that 
might be obtained if sufficient study were made of the 
plating process. Each was a reconnaissance experi¬ 
ment in the early stages of the development of the 
plating technique, designed to show whether under a 
wide variety of conditions tungsten plated by the 
vapor-phase process showed any very great differ¬ 
ences as compared with molybdenum so plated. Be- 


h They were prepared for and tested by the Bureau of 
Aeronautics, Navy Department. 


cause none of the tungsten plates was outstandingly 
better, later efforts were concentrated on molyb¬ 
denum, as described in Section 21.3. 

Before a decision can be reached concerning the 
relative value of tungsten compared with molybdenum 
plates deposited pyrolytically, a detailed investiga¬ 
tion of the factors affecting the properties of the tung¬ 
sten plates will have to be carried out, just as has 
been done with molybdenum. A possible advantage 
of tungsten plates compared with molybdenum ones 
is that harder plates of pure tungsten may be pre¬ 
pared more readily than pure molybdenum ones, and 
possibly at slightly greater plating rates. Also be¬ 
cause tungsten carbide may be somewhat harder than 
molybdenum carbide, it would seem worth while to 
continue experiments on the preparation of tungsten 
and tungsten carbide plates for use where extreme 
hardness is essential. Much work will need to be done 
before an article of irregular shape, especially one 
having deep cavities or sharp edges, can be given a 
fully satisfactory plate of tungsten or of a tungsten 
compound by this process. 

215 PREPARATION OF CHROMIUM 
CARBONYL 73 

21,51 Formation of Chromium Carbonyl 

Chromium hexacarbonyl, Cr(CO)6, the only stable 
compound formed between chromium and carbon 
monoxide, cannot be prepared in the same way as the 
corresponding molybdenum and tungsten compounds. 
It can be prepared only by use of a Grignard reagent, 
which is an organic magnesium bromide in the pres¬ 
ence of an organic solvent. The reaction, which in¬ 
volves the formation of an unstable addition com¬ 
pound, may be written as shown in equations (3) and 
(4). 

CrCl 3 + C 6 H 5 MgBr-> CrCl 3 -C 6 H 5 MgBr. (3) 

2CrCl 3 -C 6 H 5 MgBr + 12CO-> 

2Cr(CO) 6 + 2C 6 H 5 MgBr + 3C1 2 . (4) 

The exact course of the reaction and the possible 
further steps into which it may be broken down need 
not concern us here. 

The previously published work shows that several 
experimental procedures have been used in which one 
or more of the following factors were varied: the 
compound of chromium, thp type of Grignard re¬ 
agent, the solvent, and the temperature and pressure 


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435 


PREPARATION OF CHROMIUM CARBONYL 


at which the reaction was carried out. In all cases the 
yield of chromium hexacarbonyl was low, the highest 
reported being 14% of the theoretical, and duplicate 
experiments seldom gave the same yield. This is a 
common situation in organic chemistry, where more 
than one reaction is possible in a given system. The 
desired product can usually be formed in adequate 
amount only by a very careful study of the possible 
variables. 

In order to obtain better yields of the carbonyl 
than had previously been obtained, detailed studies 
were made of the following points: 

1. Chromium compound to be used as starting 
material. Chromium metal has been included. 

2. Composition of Grignard reagent. 

3. Solvent. 

4. Effect of moisture. 

5. Order and mode of addition of reagents. 

6. Type of apparatus and materials for its con¬ 
struction. 

7. Temperature at which reaction is carried out. 

8. Pressure at which reaction is carried out. 

9. Effect of stirring or other methods of agitation. 

10. Separation of the carbonyl from the reaction 
mixture. 

11. Purification of the carbonyl. 

12. Storage of the carbonyl. 

As a result of these studies, the yield of pure 
chromium hexacarbonyl was increased from about 
14% to about 64%, when prepared in quantities 
.of 5 g per run using laboratory apparatus. A few 
experiments on a slightly larger scale gave equal 
yields. No work has been done on a semicommercial 
scale, but there are some indications that a continu¬ 
ous process for the manufacture of this substance 
could be worked out without great difficulty. No 
estimate of the cost has been made. 

21.5.2 p roce( j ure f or Producing Chromium 
Carbonyl 

The procedure finally adopted was as follows: 
Commercial anhydrous chromic chloride was dried at 
250 C in nitrogen, and then screened to 40-mesh or 
finer. Moisture adsorbed here merely required the use 
of more Grignard reagent and did not seem to de¬ 
crease the yield. Phenyl magnesium bromide was 
used as the Grignard reagent. It was prepared, as 
described in standard laboratory manuals of organic 
chemistry, by dropping phenyl bromide into a flask 
containing magnesium turnings and dry ether. The 


container must have an efficient stirrer and be equipped 
with a reflux condenser. Thp reagent keeps best in an 
ether solution. 

For most of the experiments 3.08 g of chromic 
chloride were suspended in 60 cc of dry ether in a 
container that could be inserted as a tight-fitting 
liner in a high-pressure vessel. A glass container was 
used in most cases for convenience, but copper, mild 
steel, and stainless steel were also used with no 
adverse effect on the course or yield of the reaction or 
on the purity of the final product. If elevated tem¬ 
peratures are ever tried in the future, steel should be 
avoided, as unstable iron carbonyls will be formed by 
reaction with the carbon monoxide under pressure. 

The chromic chloride was kept in suspension by 
stirring in an atmosphere of nitrogen. The container 
was cooled to about —70 C, by insertion in a mixture 
of solid carbon dioxide and acetone, and then a solu¬ 
tion of 27 g of phenyl magnesium bromide in 80 cc of 
dry ether was added slowly, with continuous stirring, 
during thirty minutes. Much of the Grignard reagent 
crystallized at this temperature. The container was 
then removed from the cooling bath, its outside was 
carefully wiped, and it was placed as quickly as pos¬ 
sible in a steel pressure vessel which could be rocked 
mechanically. 

This vessel was then sealed by means of a gas-tight 
gasket and carbon monoxide was introduced through 
seamless copper or bronze tubing from a cjdinder of 
the commercial gas. Varying the pressure between 
500 and 1,000 psi did not seem to affect the yield of 
carbonyl. Pressures above 1,000 psi were not tried. 
The rocking motion was started promptly. The con¬ 
tents of the pressure vessel reached room temperature 
in about three-fourths of an hour and the reaction 
seemed to be virtually complete in 2 hr more. The 
carbon monoxide pressure was then released, the con¬ 
tents of the pressure vessel treated with ice water, 
35 cc of dilute (6N) sulfuric acid added, and the 
brownish to greenish mixture steam-distilled, as de¬ 
scribed for molybdenum carbonyl in Section 21.2.3. 

The ether was thus removed first, and the distilla¬ 
tion was continued until no more crystals of chro¬ 
mium carbonyl appeared in the receiver. Enough 
ether was then added to the receiver to dissolve all 
the carbonyl and the ether layer separated from the 
clear aqueous layer. The ether extract was washed 
with water and dried over anhydrous sodium sulfate. 
Most of the ether was distilled off through an efficient 
fractionating column. If adequate fractionation was 
not obtained, there was loss of chromium carbonyl. 


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436 


PLATING OF MOLYBDENUM, TUNGSTEN, AND CHROMIUM 


The flask containing the residue was cooled in ice, 
and the carbonyl was filtered off, washed with a little 
cold methanol and finally with a little cold dry ether, 
and dried in air at room temperature. Chromium 
carbonyl was allowed to stand exposed to the air only 
long enough for the ether to evaporate. The yield of 
crude carbonyl in these experiments was between 2.3 
and 2.9 g, that is, 53 to 67% of the theoretical. 

For further purification, the carbonyl was recrys¬ 
tallized from dry ether in a Soxhlet extractor and 
sublimed in vacuo. The sublimed material formed 
large, highly refracting crystals. Analyses for chro¬ 
mium showed the theoretical composition. 

21.5.3 Properties of Chromium Carbonyl 

Chromium carbonyl is slightly volatile at room 
temperatures and should be stored in sealed, glass- 
stoppered, dark-glass bottles. Storage in a dark cup¬ 
board is preferable, for traces of impurities make 
chromium hexacarbonyl much more sensitive to de¬ 
composition by light than molybdenum and tungsten 
carbonyls. The properties of chromium hexacarbonyl 
resemble those of the molybdenum compound very 
closely, except for the greater sensitivity to traces of 
impurities, and the slightly higher vapor pressure of 
the former. Its solubility, stability in air, and stabil¬ 
ity in contact with other materials are essentially 
those of its molybdenum analogue. Accurate values 
for the constants in question have not yet been ob¬ 
tained. 

215 4 Pilot-Plant Production 

No pilot-plant production of chromium carbonyl 
has yet been conducted. During the laboratory study 
of the synthesis, however, attention was paid to pos¬ 
sible modifications of procedure that would lead to a 
convenient and cheap commercial process. The elim¬ 
ination of two features in particular was sought—the 
use of a Grignard reagent, because of the hazard of 
working with large quantities of ether, and the use of 
high pressures, because of the complexity of the pro¬ 
cedure. No success was had in a number of attempts 
to prepare chromium carbonyl by some reaction not 
involving a Grignard reagent. 

At atmospheric pressure the yields of chromium 
hexacarbonyl by the Grignard reaction are much 
lower than when the reaction is carried out at high 
pressures of carbon monoxide. Yields of between 21 
and 24% were the best that could be obtained in 


several experiments. Starting materials other than 
anhydrous chromic chloride apparently did not react 
at all, and no carbonyl could be recovered. The yield 
of carbonyl was not increased by the use of other 
solvents or of other types ofGrignard reagent and was 
definitely reduced when preliminary temperatures 
were high. 

A few experiments involving the Grignard reaction 
under pressure on a somewhat larger scale were car¬ 
ried out toward the close of the investigation. A steel 
autoclave of 3^-gal capacity was used, and operations 
carried out at room temperature. Chromium chloride 
and ether were placed in this vessel, and the Grignard 
reagent was introduced gradually through a screw 
device, while a high pressure of carbon monoxide was 
maintained in the interior of the autoclave. The sub¬ 
sequent steps were as described above. The yield was 
64% of theory. The use of such an apparatus is sug¬ 
gested if it is desired to prepare the carbonyl on this 
larger scale. 

By the use of a specially designed apparatus, with 
a suitable number of feeding devices, a continuous 
process for the production of chromium carbonyl 
could probably be worked out if need be. The sub¬ 
sequent operations are those which have already been 
carried out on a commercial scale for many years in 
other types of organic'preparations. 

216 VAPOR-PHASE PLATING OF 
CHROMIUM 74 

2161 Introduction 

Only preliminary experiments on the formation of 
platings by the thermal decomposition of chromium 
hexacarbonyl were carried out. Enough work was 
done to show that plates of different degrees of hard¬ 
ness, which were reasonably adherent at room tem¬ 
perature, could be applied to flat steel surfaces. Tem¬ 
perature limits for the varying hardness of the 
plates were found. The few plates applied to stellite 
surfaces were not particularly adherent. 

216 2 Plating Procedure 

The process of plating chromium from the carbonyl 
was essentially the same as that of plating molyb¬ 
denum (Section 21.3). A controlled stream of hydro¬ 
gen passing over the crystalline carbonyl carried car¬ 
bonyl vapor into the plating chamber where chromium 


CONFIDENTIAL 



437 


VAPOR-PHASE PLATING OF CHROMIUM 


was deposited on the heated metal object by the 
decomposition of the carbonyl. The resulting carbon 
monoxide was removed by continuous pumping. A 
by-pass was provided in the apparatus so that the 
hydrogen could be admitted directly to the plating 
chamber when desired. 

The apparatus that was used in the chromium¬ 
plating procedure is shown to scale in Figure 10. The 
Cenco Hypervac 20 vacuum pump, which is run 
continuously during operations, is connected by a 
short piece of thick-walled rubber tubing. At the 
opposite end of the apparatus is a needle valve, 
through which a small stream of carefully purified 
hydrogen gas can be admitted. With the exception of 
these items and of the top of the plating chamber, the 
apparatus is constructed of Pyrex glass. The trap T h 
in which the unused carbonyl is condensed, is im¬ 
mersed in a bath of solid carbon dioxide and acetone. 
Traps T 2 and T z contain crystals of purified chromium 
hexacarbonyl; T 2 is surrounded by a constant-tem¬ 
perature bath (not shown), while T z is at room tem¬ 
perature. The steel or other disk to be plated is 
suspended in chamber C by means of the thermo¬ 
couple leads which are sealed into a ground glass 
stopper set into the heavy brass top of the chamber. 
The chamber is surrounded and heated by an induc¬ 
tion coil (not shown). Also shown are connections to 
a McLeod gauge to indicate pressure, to a standard 
12-liter volume for calibration of gas flow, and at S z , 
to a source of hydrogen sulfide or other gas, which 
may be used in addition to the hydrogen. 

In this apparatus the disk temperature can be 
varied from room temperature to about 1000 C and 


carbonyl vapor pressure from 0.040 to 0.220 mm. The 
hydrogen flow and the proportion of any other gas to 
hydrogen may be varied at will. The vapor pressure 
of the carbonyl can be calculated from the tempera¬ 
ture of the bath surrounding the carbonyl container 
T 2 by equation (5), 

log p = 10.63 - 3285/T (5) 

in which p is expressed in mm of Hg and T in degrees 
centigrade, provided that the bath has been in posi¬ 
tion for sufficient time for temperature equilibrium 
to have been established with the carbonyl. 

The procedure preliminary to plating was the same 
as for molybdenum plating in that the metal surfaces 
had to be deoxidized by heating in hydrogen at a 
temperature higher than the plating temperature. 
When the specimen had cooled to the desired plating 
temperature, the temperature of the carbonyl supply 
was adjusted to provide the desired vapor pressure, 
and the vapor was admitted to the system. After 2 
to 4 hr, the carbonyl vapor was cut off, the sample 
was again heated to the deoxidizing temperature (700 
to 800 C), and cooled gradually to anneal the plate. 

21,6-3 Characteristics of the Plates with 

Variations in Procedure 

Because the high temperatures involved in this 
procedure permanently change the characteristics of 
the steel, some runs were made in which the tempera¬ 
ture was never raised above 550 C (well below the 
alpha-gamma transition point of steel). This resulted 
in a slight decrease in the strength of the bond be- 



Figure 10. Apparatus used in the experiments on the vapor-phase plating of chromium. (Figure 1 in NDRC Report 
A-402.) 


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438 


PLATING OF MOLYBDENUM, TUNGSTEN, AND CHROMIUM 


tween steel and plate and also involved a longer time 
to yield the same plate thickness. 

Plates on stellite disks showed poor adherence 
when the standard procedure was used. This was 
presumably due to incomplete deoxidation of the 
stellite surface before plating. Chromium oxide is less 
readily reduced than is iron oxide, and the experi¬ 
ments with molybdenum plating showed that a good 
stellite-to-plate bond requires deoxidation for x /i hr 
at 1100 C. Experiments to confirm this with chromi¬ 
um plates were not carried out. 

The composition and hardness of the plate appear 
to be determined primarily by the disk temperature 
during the plating process. The results of about 40 
experiments on steel disks are summarized in Table 5. 


Table 5. Composition and hardness of pyrolytically 
deposited chromium plates. 


Plating 

temperature 

(C) 

Approximate 

composition 

(%) 

Approximate 

hardness 

(Vickers) 

Quality of 
disk-to-plate 
bond 

250-350 

50 

Ci ’203 

1400 

Very poor 


50 

O3C2 



400-500 

95 

Cr 

1200 

Good 


5 

O2O3 



500-525 

50 

Cr 

2000 

Poor 


25 

Cr 2 0 3 




25 

Unknown 



530-550 

50 

Cr 


Very poor 


50 

Cr 2 0 3 



600-650 

40 

Cr 

Very hard 

Good 


60 

Cr 2 0 3 

(2000) 




4- Cr 3 C2 




Identification of the compounds in the plate was by 
x-ray analysis. The unknown constituent may be a 
hitherto unknown carbide of chromium. 



At plating temperatures of about 400 C a well- 
bonded, comparatively soft plate is formed, and at 
about 625 C a very much harder plate with a good 
bond is formed. A chromium oxide-chromium carbide 
plate appears to possess poor adherence, while most 
of the plates carrying metallic chromium appear to 
adhere better. The presence of oxide in these plates, 
in contradistinction to the absence of oxides of mo¬ 
lybdenum or tungsten in plates of those metals, 
should be noted. This is presumably due to the much 
greater ease of formation of chromium oxide and to 
its greater stability after having been formed. 

The plates prepared varied from 0.5 to 5.7 mils in 
thickness. No relation between thickness and adher¬ 
ence could be noted because of other variables. The 
plating rate was approximately proportional to the 
vapor pressure of the carbonyl and, at low rates of 
flow, to the hydrogen flow. At higher rates of hydro¬ 
gen flow, the gas was probably not saturated with 
carbonyl vapor. The disk temperature did not seem 
to affect the plating rate except at the lowest temper¬ 
atures. The presence of hydrogen sulfide, added in 
very small amounts in an attempt to deoxidize the 
disk surface, decreased the plating rate enormously 
and ruined the bond. The plate deposited under such 
conditions is 95% pure chromium with a hardness of 
only 350 Vickers. 

Photomicrographs of four typical plates are shown 
in Figure 11. In all cases the magnification is 500X. 
The first etchant used was 5% nital, to show the 
structure of the steel, followed by a boiling solution 
of potassium hydroxide and potassium ferricyanide, 
to show the structure of the plate. The top of each 
section shows the electroplated surface of iron applied 
to protect the specimen during cutting and polishing. 



Figure 11. Photomicrographs of transverse sections of chromium plates deposited on steel disks by vapor-phase deposi¬ 
tion. The structure of the plates varied considerably depending on the temperature of the specimen. ( 500X .) (Portions of 
Figures 2 and 3 of NDRC Report A-402.) 


CONFIDENTIAL 









VAPOR-PHASE PLATING OF CHROMIUM 


439 


216 4 Conclusions and Recommendations 

No attempts have been made to apply chromium 
pyrolytic plates to gun bores or to the inside of smooth 
tubes. Experience with molybdenum pyrolytic plates 
makes it appear doubtful if success could be obtained 
with the expenditure of only a reasonable amount of 
effort. Moreover, because of the development of al¬ 
loys (Chapters 17 and 18) that withstand tempera¬ 
tures at which stellite fails, further investigation of 
chromium pyrolytic plates as a means of extending 
the usefulness of stellite to hypervelocity guns is not 
considered worthwhile. 

These plates of chromium-chromium oxide-chro¬ 
mium carbide on steel, however, are of considerable 
general interest. We have here another type of plate 


that can be applied in an adherent form to a metal 
surface, the hardness of wdiich can be varied 
nearly at will throughout remarkably wide limits, 
and w r hich is (though this needs confirmatory tests) 
very resistant to corrosion. 

Further study should be made of the conditions 
under w r hich an adherent plate of given properties 
can be deposited at a reasonable rate. There should 
also be further study of the firmness of the bond, and 
of methods to improve it. These plates have promise, 
it would appear, in the field of extremely hard 
protective coatings for small articles; also for precision 
bearing surfaces, or for surfaces subjected to hard 
wear. The present stage is obviously only a very 
preliminary one, and extensive research must be done 
before determining the limits of their application. 


CONFIDENTIAL 










PART VI 

IMPROVED MACHINE GUN BARRELS 



A hit, a very palpable hit. 
—William Shakespeare 
“Hamlet” 


CONFIDENTIAL 
































































































* 










* 























Chapter 22 

STELLITE-LINED MACHINE GUN BARRELS 

By J. F. Schairer a 


221 INTRODUCTION 

A short breech LINER of an erosion resistant ma¬ 
terial in a gun barrel, which provides a suitable 
bore surface at and near the origin of rifling where 
erosion is most severe, is a practical method of de¬ 
creasing erosion and thus improving gun barrel per¬ 
formance. In order to utilize this method of mitigat¬ 
ing erosion in machine gun barrels it was necessary to 
find and develop a suitable liner material and to in¬ 
sert it in a practical way in a Service weapon. Divi¬ 
sion 1 has demonstrated that such an erosion-resist¬ 
ant liner material existed, that it could be inserted in 
a practical manner in a Service weapon, that it would 
remain in place during firing with no rotation, for¬ 
ward or backward movement, or serious expansion or 
contraction, and that a gun barrel with such a liner 
(with the liner joint and discontinuity in the bore sur¬ 
face) would fire accurately and safely and show such 
outstanding improvement in life and performance as 
to justify amply the expense and effort required for 
the modification of regular steel barrels. 

The stellites, because of their resistance to chem¬ 
ical attack by hot powder gases, hardness and strength 
at high temperatures, and excellent wear and abra¬ 
sion resistance, have shown outstanding performance 
in machine gun barrels under both mild and severe 
firing conditions, including schedules so severe that 
unmodified steel barrels are unable to withstand 
them. The experiments that led to the selection of 
Stellite No. 21 as an eminently suitable liner material 
for machine gun barrels have already been described 
in Chapters 16 and 19. The development of a prac¬ 
tical design for liner insertion in machine gun barrels 
for production and Service use is described in the 
present chapter. 

22 2 DEVELOPMENT OF THE LINER DESIGN 
22,21 Insertion of Short Breech Liners 

Early in the erosion-resistant materials program of 
Division 1, NDRC (Section 16.1), the importance of 

a Special Assistant, Division 1, NDRC. (Present address: 
Geophysical Laboratory, Carnegie Institution of Washington.) 


liner insertion for the testing of special materials was 
realized. Recognizing that the only conclusive test of 
• an erosion resistant material is a firing test in the gun 
in which it is to be used or under conditions as close 
as possible to this ultimate objective, the several 
types of experimental liners described in Sections 16.3 
and 26.2 were devised. 

All of these designs for liner insertion were devel¬ 
oped in order to test liner materials and not overall 
barrel performance. None of them was intended to be 
suitable for mass production or for Service use. 

In order that an erosion resistant material, once 
found, could be applied, the Crane Company was 
asked in August, 1943 to develop a method of insert¬ 
ing short breech liners that would be practical from 
both the production and Service standpoints. In view 
of the urgent need for improvement in barrel life and 
performance of the caliber .50 Browning machine 
gun, particularly for aircraft combat, all efforts were 
concentrated initially on barrels of this caliber. Fur¬ 
thermore, it was considered desirable initially to main¬ 
tain the same external barrel contour in order to 
utilize guns, mounts, and installations already in 
Service use and to develop a method for the modifica¬ 
tion of regular steel barrels already in large supply to 
permit liner insertion. 

Crane Company prepared four designs for liner in¬ 
sertion in caliber .50 machine gun barrels, two for the 
heavy and two for the aircraft barrel. Three of the de¬ 
signs were similar in that the assembly consisted of 
three parts—a liner, a chamber section, and a for¬ 
ward section. The liner, which was tapered, was 
pressed and shrunk into the chamber section and the 
forward section was then pressed on and the assembly 
held together with tapered pins. In the fourth design, 
applied to an aircraft barrel, the liner and chamber 
sections were integral. 

Firing tests were conducted on one heavy barrel 
assembly made in accordance with what was consid¬ 
ered to be the most promising of these designs, using 
a steel liner for test purposes. After only a few rounds 
had been fired, the tapered pins sheared and the sec¬ 
tions of the barrel separated. Subsequently, another 
design, similar to this except that a union-ring type 
of nut rather than tapered pins was used and was 


CONFIDENTIAL 


443 



444 


STELLITE-LINED MACHINE GUN BARRELS 



Figure 1. Stellite liner 9-inches long inserted in caliber .50 aircraft machine gun barrel according to Crane Company 
drawing No. 50-35. (Figure 2 of NDRC Report A-408.) 


tested with a steel liner with satisfactory results. 
However, this design was not completely satisfactory 
as far as production was concerned and could not 
readily be adapted to the aircraft barrel for which 
there was the most urgency for improvement. 

A new design for the aircraft barrel was developed 
in which a tapered stellite liner was.inserted from the 
breech end of the barrel and held in place by a thread¬ 
ed retainer screwed into a tapped recess between the 
breech ends of the liner and barrel. During firing, the 
tapered stellite liner moved forward, upsetting the 
steel ahead of the liner and constricting the steel bore. 
This forward movement occurred as a result of the 
low coefficient of friction between stellite and steel 
and dimensional changes in the liner caused by 
stresses during firing. Increasing the taper or rough¬ 
ening the liner surface by sandblasting were ineffec¬ 
tive in preventing the forward movement of the liner. 

As a result of these difficulties, the design was 



Figure 2. Caliber .50 stellite-lined aircraft machine 
gun barrel disassembled to show barrel, liner, and 
retainer. A complete round of ammunition is also 
shown. (Figure 3 of NDRC Report A-408.) 


further modified to provide for a cylindrical liner with 
an enlarged shoulder or flange at the breech end to 
prevent forward movement. The liner was inserted 
on a diametral shrink-fit by heating a recessed barrel 
to about 450 or 500 F and cooling the rifled liner in 
liquid air or nitrogen. Rifling in the liner and barrel 
were put in alignment during insertion by the use of a 
rifled mandrel. This design which is shown in Figure 
1, entirely eliminated the forward movement of the 
liner and proved entirely satisfactory and practical 
for production. The stellite-lined barrel was adopted 
for Service use in both the caliber .50 aircraft and 
heavy barrels as described in Sections 22.3.4 and 
22.4.2. A stellite-lined caliber .50 aircraft machine 
gun barrel, disassembled to show the recessed barrel, 
liner, retaining nut and cartridge, is shown in Figure 2. 


22 2 2 Firing Tests on Barrels with Liners 
of Stellite No. 21 


First Firing Test 


For the testing of various liner materials to single 
out those which are markedly superior to gun steel in 
overall erosion resistance, it was desirable to use a 
severe firing schedule. Because the heavy walls of the 
caliber .50 heavy machine gun barrel gave good sup¬ 
port to a liner, this barrel was used in initial liner 
tests. The following severe schedule was used for the 
testing of the first liner of Stellite No. 21: 


Preliminary 
rounds 
(total 145) 


1 single round 

9 single rounds in succession 
1 burst of 10 rounds 
1 burst of 25 rounds 
1 burst of 100 rounds 


500-round 

group 


5 bursts of 100 rounds each with 1 
min between beginning of each 
burst. 


The 500-round groups were to be repeated as many 
times as necessary. The barrel was cooled to room 


CONFIDENTIAL 







































DEVELOPMENT OF THE LINER DESIGN 


445 


temperature, examined, and gauged with a breech- 
bore gauge after each item of the preliminary schedule 
and after each 500-round group. 

The first liner of Stellite No. 21 tested was an in¬ 
vestment-cast liner, 6 in. in overall length. This cyl¬ 
indrical flanged liner was inserted in a caliber .50 
heavy machine gun barrel according to a design es¬ 
sentially the same as that shown in Figure 1. After 
over 6,000 rounds had been fired, the breech-bore 
gauge had advanced only 0.30 in. (A regular steel bar¬ 
rel is declared worn out after an advance of 2 in. of 
this gauge.) The liner was still in good condition, but 
since the liner was too short, the steel bore surface 
ahead of the liner had eroded considerably. The liner 
was removed for very careful examination, inserted in 
a new barrel, and the firing continued. At 10,000 
rounds, the breech-bore gauge had not advanced as 
much as 1 in.; but shortly thereafter the advance be¬ 
came more rapid and the rejection point (2 in.) was 
reached at 10,900 rounds. The liner b had not failed in 
a dangerous manner by cracking but gradually wore 
out. Such liner performance was phenomenal. On the 
same firing schedule a steel barrel was worn out in 
less than 1,000 rounds. 

Firing Tests on Aircraft Barrels 

Steps were taken immediately to apply a liner of 
Stellite No. 21 to the caliber .50 aircraft machine gun 
barrel and to learn the overall improvement in barrel 
performance that could be obtained by the use of such 
a liner. No measurements were available on the ef¬ 
fects on accuracy and velocity life of a gun barrel 
when a short breech liner of an erosion resistant ma¬ 
terial had been inserted. It was not possible to get 
such measurements until a suitable liner material had 
been found and satisfactorily inserted. Division 1 
made arrangements with the Sixth Service Command 
for the Crane Company to utilize an Army firing 
range at Fort Sheridan near Chicago to make accu¬ 
racy and velocity firings on stellite-lined barrels. 

Two stellite-lined caliber .50 aircraft barrels were 
to be prepared and compared with two standard bar¬ 
rels on the same moderately severe firing schedule. In 
the preparation of the lined barrels the liner length 
was increased to 9 in. This length was a compromise. 
The firing test in the heavy barrel described above 
had shown that a 6-in. liner was too short. Owing to 


b Affectionately called “Grandpappy” by those associated 
with the test. 


practical matters, such as maintenance of tolerance 
during the recessing of barrels to receive liners and 
the limitations imposed by the wall thickness of the 
light aircraft barrel, it was decided to determine the 
performance of the 9-in. liner. 

While the 9-in. investment-cast liners were being 
prepared, while tooling was in progress for their in¬ 
sertion, and while equipment was being assembled for 
the accuracy and velocity firings, a composite 73^-in. 


9 —• 



L » I , I , I . I 

/ 2 3 4 


Figure 3. Target sheet for 25-round burst for cold 
accuracy at 1000 inches during firing test of aircraft 
barrel No. 5 (fused Stellite No. 21 liner), after 5,084 
rounds. Target scale in inches. (Photograph No. 70 in 
Crane Monthly Report on Contract OEMsr-629 for 
June 1944.) 


CONFIDENTIAL 





446 


STELLITE-LINED MACHINE GUN BARRELS 


liner was prepared by casting Stellite No. 21 into a 
steel sleeve. This liner was inserted in an aircraft bar¬ 
rel and subjected to a firing test. In the light aircraft 
barrel the same firing schedule as used on the heavy 
barrel was unsatisfactory. A preliminary test on a 
steel barrel showed that such a schedule was too 
severe and that the barrel failed by softening of the 
steel barrel wall near the muzzle. The schedule adopt¬ 
ed for initial tests with lined aircraft barrels, con¬ 
sisted of 250-round groups, with complete cooling be¬ 
tween groups, each group consisting of 50-round 
bursts 1 minute apart. In the firing test (Crane Barrel 



Figure 4. Moving target and mechanism for record¬ 
ing burst fire. (Photograph No. 100 in Crane Monthly 
Report on Contract OEMsr-629 for August 1944.) 


No. AC5) on the composite 7-in. liner, the breech- 
bore gauge entered only 0.35 in. after 5,000 rounds 
on this schedule and the rifling in the liner was still 
sharp, and the liner bore surface showed only a few 
small pits and hairline surface cracks. The cold accu¬ 
racy stationary target sheet after this test is shown in 
Figure 3. 

The accuracy and velocity firings of four caliber .50 
aircraft barrels, two with flanged 9-in. investment- 
cast liners or Stellite No. 21 inserted on a shrink-fit 
(Figure i) and two standard steel barrels for compar¬ 
ison, showed the outstanding overall barrel perform¬ 
ance of stellite-lined barrels. The firings were con¬ 
ducted with standard ammunition, AP-M2. The ini¬ 
tial muzzle velocity of the barrels was about 2,700 


fps. In the case of the lined barrels after twelve 250- 
round groups (3,000 “erosion rounds”) no significant 
velocity drop either hot or cold had occurred. With 
the standard steel barrels, velocity had decreased to 
2,100 fps while hot and 2,500 fps while cold after only 
two 250-round groups (500 “erosion rounds”)- As 
shown by a moving target (see Figure 4) there was no 
inaccuracy (keyholing bullets) during the 3,000 ero¬ 
sion rounds with the stellite-lined barrels and the dis¬ 
persion was not significantly less than the initial 
value. On the other hand, with the standard steel bar¬ 
rels after only two erosion groups (500 rounds) the 
bullets covered a target area 40 ft in diameter at 300 
yd and all rounds keyholed during the fourth and 
fifth bursts of the second 250-round group. The firing 
test on the two stellite-lined barrels was arbitrarily 
discontinued after 3,000 erosion rounds. This firing 
schedule was not severe enough to show the full po¬ 
tentialities of these barrels. 


22 3 APPLICATION OF STELLITE NO. 21 
TO THE CALIBER .50 AIRCRAFT 
BARREL 

22 3 1 Need for Improvement 

Experience in aircraft combat by both the Army 
and Navy Air Forces and those of our British allies had 
indicated that barrel life and performance of the reg¬ 
ular caliber .50 aircraft machine gun barrel were in¬ 
adequate to meet the needs and requirements of severe 
combat use. Because of the speed and maneuvera¬ 
bility of the target, a high rate of fire and high velocity 
were required in plane-to-plane combat. Movement 
of the high-speed target during the time-of-flight of 
the bullets required the shortest possible time-of- 
flight. Velocity drop destroyed the accuracy of the 
computing sights used. The necessity for hitting a 
target many times before it was destroyed required 
high velocity, high rate of fire, and longer bursts of 
fire. The advent of long-range bombing and strafing 
attacks and night fighting astern of a target located 
by radar caused a great increase in the length of 
bursts desired in machine guns. The regular caliber 
.50 aircraft barrel did not stand up under this abuse 
and the ability to fire long bursts became an urgent re¬ 
quirement for this weapon. 

Barrel erosion was seriously limiting the life and 
performance of these machine gun barrels. If barrel 
erosion could be mitigated, vastly more efficient 


CONFIDENTIAL 








STELLITE APPLIED TO CALIBER .50 AIRCRAFT BARREL 


447 


weapons could be placed in the hands of our fighting 
forces. On the basis of the firing tests on the stellite- 
lined barrels just described, the utilization of the re¬ 
markable erosion resistance of Stellite No. 21 when 
applied as a short breech liner, showed great promise 
of yielding a vast improvement in the overall per¬ 
formance and life of this important and critical Serv¬ 
ice weapon. 

22,3,2 Preparation of Lined Barrels for 
Ordnance Department Tests 

As soon as the accuracy and velocity firings on the 
two stellite-lined caliber .50 aircraft machine gun bar¬ 
rels had proceeded far enough to confirm the expecta¬ 
tion of the remarkable increase in overall barrel per¬ 
formance to be obtained by the use of a short erosion- 
resistant breech liner, Division 1 made plans for the 
immediate application of these results. Although it 
was realized that the design and perhaps even the 
exact liner material might not be as suitable as that 
which might be revealed by further research, it was 
felt that the improvement in performance was so out¬ 
standing that it would be desirable to freeze the 
material and the design around that shown in Crane 
Drawing No. 50-35 (already shown as Figure 1) and 
furnish barrels for Ordnance Department acceptance 
tests and immediate Service application, rather than 
wait for possible further improvements. 

Accordingly Division 1 asked Crane Company to 
prepare 212 barrels just like those which had shown 
such superior performance in the accuracy and veloc¬ 
ity firings. Two hundred of these barrels were for 
test by the Small Arms Division, Research and De¬ 
velopment Service of the Ordnance Department and 
12 by the British Air Commission. Crane Company 
prepared these barrels under Contract OEMsr-629 by 
the modification of standard steel barrels supplied by 
the Ordnance Department.The investment-cast liners 
were supplied by the Haynes Stellite Company (sub¬ 
sidiary of Union Carbide and Carbon Corp.) under 
Contract OEMsr-1330. With the special tools on 
hand at Crane Company, it was possible to prepare 
two barrels per day and deliver them for immediate 
test. Additional tooling was expedited and soon four 
barrels per day were available. Because of the long 
life and superior performance of the stellite-lined bar¬ 
rels, it was possible to supply these barrels faster than 
they could be tested. In accordance with directions 
from the Ordnance Department, 150 barrels were 
sent for test to the Army small arms testing range at 


Purdue University and 50 to Aberdeen Proving 
Ground. All of these barrels were made and delivered 
by October 30, 1944. 

The only dimension changed during the course of 
the preparation of these barrels was that of the bullet 
seat whose diameter was increased by 0.002 in. This 
cylinder (32.1 in. from the muzzle) was made slightly 
larger in stellite-lined c barrels than in standard steel 
barrels because of the slight bore constriction (about 
0.002 in. in diameter) as a result of firing. If the bul¬ 
let seat is too small, stoppages result because of fail¬ 
ure of the cartridge to seat. In standard steel barrels, 
erosion more than keeps pace with constriction. 

22,3,3 Test Results 

In order to evaluate the relative life and perform¬ 
ance of the stellite-lined caliber .50 aircraft machine 
gun barrels prepared by Division 1, as compared with 
standard steel barrels under the same firing condi¬ 
tions, the Small Arms Division, Research and Devel¬ 
opment Service of the Ordnance Department, utilized 
a series of firing schedules which gave a measure of 
relative barrel performance under conditions simulat¬ 
ing those that might be used in combat. They varied 
from the very severe conditions of long continuous 
bursts or moderate length bursts (100 rounds) at fre¬ 
quent intervals (2 min apart) to such mild schedules 
as 100-round bursts with complete cooling between 
bursts. The effects on performance were also eval¬ 
uated for the various types of caliber .50 ammunition, 
including ball, M2; armor-piercing, M2; armor-pierc¬ 
ing incendiary, M8 and M8E1; incendiary, Ml; 
tracer, Ml and M10; and other special experimental 
ammunition. 

Under all conditions, but particularly with long 
continuous bursts of fire so much desired in severe 
combat, the stellite-lined barrels showed a very 
marked superiority to the standard steel barrels. Al¬ 
though the stellite-lined barrels showed a marked im¬ 
provement in accuracy life over standard steel bar¬ 
rels, the improvement in velocity life was phenomenal. 
Owing to a small permanent bore contraction in stel¬ 
lite-lined barrels during the first burst of fire, the 
velocity actually increased and then decreased only 
very slowly during additional firing. In the case of the 
tests illustrated by Figure 8, the initial increase 
amounted to 150 fps. 


c Also in nitrided and chromium-plated barrels described in 
Chapter 23. 


CONFIDENTIAL 





448 


STELLITE-LINED MACHINE GUN BARRELS 


/ . ..... 

^10 

30 

\ "" 

20 

40 50 

\ • / 

' 1 # 

* * , , » . • 

. • * * 

•X 

; Vs 

SO 

.•. • / 

% H 

90 

• / , ' . 

V 

100 

HO 120 

. / ' 

/* 

140 150 

\, 

A • * , N ** • 

170 , 

T . > / 

' .>■% 

■ ' N i 

210 

4 

/ 

200 

' i' ■ • 

. 220 


250 

' • ■ 7 

260 

270 

/ 

> 

280 290 

v / 

% 

/ 

300 

310 320 

i ..>•. 

» 

330 

3-MISSED tarcet 

OR TIRED THRU 

. SAME HOLES 

T . 350 

347 




RNDS.39-388 

EROSION TEST N2] 

MOVING TARGET AT 1000" 

CAL. .50 AIRCRAFT BARREL AC42 

0 6' 12* 18*^“ 

^ * . ... CRANE CO 

1-15-45 

Figure 5. Moving ta 
(Photograph No. 283 ir 

rget pattern at 1,000 in. for first burst (350 rounds) in test of stellite 
i Crane Monthly Report on Contract OEMsr-629 for December 1944.) 

-lined barrel No. AC42. 

, 

20 

, i 

% • 

“ V 

30 

40 50 

! \ 

§ 

. * 60 « 
v 

1 « f \ 1 
l 80 ,4- 

JAM 

90 3.5SEC 

v- ■ -1 

JAM 

100 1 SEC 

\ V 

I 

• , *, 

" \ 

110 

"j. 

120 

N * * 

> ' / 

130 

140 

7 - * V •' 
/ , 160 

- 171 

. / 

X 

170 



RNDS.2059-2229 

EROSION TEST N2 8 

MOVING TARGET AT 1000" 

CAL. .50 AIRCRAFT BARREL AC42 

0 6" 12" 18" @ 

1.■.1 . ■ . . ,j CRANE CC 

1-15-45 


Figure 6. Moving target pattern at 1,000 in. for eighth burst (171 rounds) in test of stellite-lined barrel No. AC42. 
(Photograph No. 338 in Crane Monthly Report on Contract OEMsr-629 for December 1944.) 


CONFIDENTIAL 




























AVERAGE INSTRUMENTAL VELOCITY AT 78 FT IN FPS 


STELLITE APPLIED TO CALIBER .50 AIRCRAFT BARREL 


449 


BEFORE EROSION 

AFTER EROSION 

• 

NCL1 _ 

, N0.2 



••v 


£ . 

V- 

• 



3-27 

• • 

393-417 

641-665 


NO. 6 

I . 

V* 


1818-1842 


NO. 7 

I • 




2034-2058 


AFTER EROSION 
NO. 8 


I'f 


2245-2269 


AFTER EROSION 
N0.3 ■ NO.4 « 




982-1006 




1255-1279 


AFTER EROSION 

( 

NO, 5 ■ NO. 5 

^ I V 


1495*1519 


1520-1544 


CAL .50 AIRCRAFT BARREL AC42 
COLD ACCURACY AT 1000" 

25 RND. BURSTS 

6 ^?CRANE CO. 

‘ 1 1 ' 1 . 1 1-15-45 


0 


Figure 7. Target sheets for 25-round bursts for cold accuracy at 1,000 in. during the course of the firing test on stellite- 
lined barrel No. AC42. (Photograph No. 348 in Crane Monthly Report on Contract OEMsr-629 for December 1944.) 



Figure 8. Velocity loss in standard steel and stellite-lined caliber .50 aircraft machine gun barrels on the 100-round burst 
cycle (F-l). (Graph prepared from data in Aberdeen Proving Ground Third Memorandum Report on O.P. 5082.) 


CONFIDENTIAL 


MEAN VARIATION IN FPS 





















450 


STELLITE-LINED MACHINE GUN BARRELS 


The improvement in performance on various test 
schedules d may be summarized as follows: 

Continuous Burst Cycle (C — l ). In this very severe 
schedule the barrel is fired in a continuous burst until 
serious keyholing® of the projectile develops. “Cold- 
velocity” is measured before and after the burst and 
sometimes “hot-velocity” is measured at the end of the 
burst. Continuous bursts to serious keyholing are re¬ 
peated (with complete cooling between bursts) until a 
drop in cold-velocity of 200 fps occurs and the barrel is 
rejected for excessive velocity drop. Results on this 
test schedule show the remarkable superiority of the 
stellite-lined barrel for use in strafing and by night 
fighters. 

Ordnance Department tests on the steel barrels 
that were standard in 1943 (Drawing No. D-28272) 
showed that after a single burst of 167 rounds (mean 
of ten barrels) nearly complete keyholing occurred 
and there had been a velocity drop of nearly 200 fps. 
Thus life was only about 167 rounds. Similar tests 
were made by the Ordnance Department on seven 
stellite-lined barrels provided by Division 1. These 
seven barrels averaged 295 rounds to serious keyhol¬ 
ing for the single continuous burst that was fired. 294 
No velocity measurements were made on these barrels. 

Division 1 conducted more extended tests on reg¬ 
ular stellite-lined barrels on the continuous schedule 
at Crane Company to obtain a standard for compar¬ 
ison with other special barrels being studied in the 
research program. AP-M2 ammunition was used. Al¬ 
though there was some variation in life and perform¬ 
ance of regular 1 stellite-lined barrels, the firing results 
on barrel No. AC42 80 are representative of what 
performance was achieved on this schedule. Eight 
continuous bursts of 350, 212, 301, 233, 200, 247, 176, 
and 171 rounds, respectively, were fired, plus 379 in¬ 
termittent rounds for accuracy and velocity measure¬ 
ments, making a total of 2,269 rounds. The cold- 
velocity dropped only 15 fps after the first burst of 

d Some of these same firing schedules were used in the test¬ 
ing of nitrided and chromium-plated caliber .50 barrels, as de¬ 
scribed in Sections 23.1.3 and 23.1.4. Stellite-lined barrels were 
later tested on the CGL-350 schedule. (See Sections 24.1.3 
and 24.4.) 

e Inaccurate (tumbling) bullets hit the target paper broad¬ 
side-on (British “B.S.O.”) and produce a keyhole-shaped hole 
in the target paper instead of a round hole. 

f A “regular” stellite-lined barrel in this report means a 
caliber .50 steel barrel made to Ordnance Drawing No. D-28272 
in which a 9-in. liner (Figure 1) of investment-cast Stellite No. 
21 has been inserted by the method developed by the Crane 
Company 80 (described at the end of Section 22.2.1) or by some 
equivalent method. 


350 rounds. After that it decreased only gradually 
from its initial value of 2,705 fps to 2,515 fps after the 
eighth burst. The moving target patterns during 
the first and eighth bursts are shown in Figures 5 
and 6, respectively. The stationary targets for 25- 
round bursts for cold accuracy at 1,000 in. during the 
course of the test with this barrel are shown in Figure 
7. This stellite-lined barrel not only showed a life of 
1,890 rounds as compared with 167 for a standard 
steel barrel, but in addition it fired an initial burst of 
twice the length.® Even the eighth burst was equal in 
length to the first and only burst possible with the 
standard steel barrel. 

100-Round Burst Cycle (B — l ). In this moderately 
severe schedule the barrel is fired in 100-round bursts 
with two-minute cooling intervals until serious key¬ 
holing develops or the velocity drop is greater than 
200 fps. Ordnance Department tests on 11 standard 
steel barrels fired with combat ammunition showed 
an accuracy life of 230 rounds while similar tests on 
13 stellite-lined barrels showed an accuracy life of 455 
rounds. No velocity measurements were made on the 
standard steel barrels but the stellite-lined barrels 
showed a velocity increase of 30 fps during the firing. 
When seven standard steel barrels were fired, with 
ball M2 ammunition for five 100-round bursts with 
2-min cooling intervals, the velocity drop was 533 fps. 

100-Round Burst Cycle ( F-l ). In this mild schedule 
the barrel is fired in 100-round bursts with complete 
cooling between bursts. Velocity and accuracy tar¬ 
gets are taken during every fifth burst and the cycle 
is repeated until a 200-fps drop in cold-velocity oc¬ 
curs, or a majority of the rounds fired in a burst 
produce keyholes. 

Three stellite-lined barrels and three steel barrels 
were fired at Aberdeen Proving Ground on this sched¬ 
ule using combat (API-M8) ammunition. 213 Velocity 
drop was the cause of end of life in both cases. The 
standard steel barrels showed a life of 2,500 rounds 
while the stellite-lined barrels showed a life of 7,000 
rounds. Even on this mild schedule stellite-lined bar¬ 
rels showed nearly three times the life of a standard 
steel barrel. The velocity loss data for these tests is 
shown graphically as Figure 8. 

300-Round Burst Cycle (.British Schedule). In this 
severe schedule the barrel is fired in 300-round bursts 
with complete cooling between bursts. Cold-velocity 
measurements are made at the start of the test and 

B In tests of a large number of barrels at Crane Company the 
length of the initial continuous burst to serious keyholing 
varied between 290 rounds and 350 rounds. 


CONFIDENTIAL 





STELLITE APPLIED TO CALIBER .50 AIRCRAFT BARREL 


451 


after each burst. Accuracy is determined throughout 
a burst by means of a moving target. Breakdown of 
accuracy is taken as the stage at which frequent 
tipping bullets and partial B.S.O.’s (keyholing bullets) 
commence. 

A portion of the stellite-lined barrels prepared by 
Division 1 for the British Air Commission were tested 
on this schedule. The results of test 406,413 on two 
barrels with AP-M2 ammunition showed an accuracy 
life of 850 and 880 rounds, respectively, as compared 
with 170 rounds (mean of ten barrels) on standard 
steel barrels. Firing was continued on the stellite- 
lined barrels and a 200-fps velocity drop occurred 
finally after 1,120 and 1,160 rounds, respectively, as 
compared to about 180 rounds in the standard steel 
barrels. 

Ten 40 -Round, Bursts (British Schedule). In this 
test schedule which the British used to indicate suit¬ 
ability of a barrel for improved performance in fighter 
combat use, ten 40-round bursts are fired at the be¬ 
ginning of each minute. The cycle is repeated following 
complete cooling after each 400-round group. 

A portion of the stellite-lined barrels prepared by 
Division 1 for the British Air Commission were tested 
on this schedule. 406 413 The results of test on two 
barrels with AP-M2 ammunition showed an accuracy 
life of 3,560 and 4,330 rounds, respectively, as com¬ 
pared with only 190 rounds for the standard steel 
barrel. Firing was continued on the stellite-lined bar¬ 
rels and a 200-fps velocity drop occurred finally after 
4,300 and 5,200 rounds, respectively, as compared to 
a 400-fps velocity drop in only 240 rounds with a 
standard steel barrel. Such a marked superiority of 
roughly 20 to 1 of stellite-lined barrels over standard 
steel barrels in a test designed to indicate the suit¬ 
ability of a barrel for combat use in fighter planes 
showed convincingly the remarkable potentialities of 
the stellite-lined barrel. 

22.3.4 Adoption for Service Use 

As a result of the very favorable acceptance test 
results on the 200 stellite-lined barrels prepared by 
Division 1 on the firing schedules just described, the 
stellite-lined caliber .50 aircraft machine gun w T as 
adopted as standard for Service use. At a meeting of 
the Ordnance Committee 294 on January 4, 1945 ap¬ 
proval was given to recommendations that “Barrel 
(D-7161580) having a stellite liner be approved for 
manufacture as the preferred design for use in Gun, 
Machine, Browning, Caliber .50, M2, Aircraft, Basic ; 


Gun, Machine, Caliber .50, T36, Aircraft, Basic; and 
Gun, Machine, Caliber .50, T25E3, Aircraft, Basic” 
and that “The Barrel Assembly presently in manu¬ 
facture 11 (D-28272) be produced only in such quantity 
as may be required to balance the Army Supply 
Program, for the Gun, Machine, Caliber .50, M2, 
Aircraft, Basic; provided that production of Barrel 
Assembly (D-28272) shall terminate at such time as 
requirements can be met by the preferred types.” 1 

22.3.5 pi} ot pi an t Production of Barrels for 
Extended Service Tests in Combat 

The first accuracy and velocity firings on stellite- 
lined barrels by Division 1 and all subsequent tests 
by the Ordnance Department 319 were so favorable 
that it was imperative to develop efficient means of 
manufacture and to prepare barrels for extended Ser¬ 
vice tests in combat areas at the earliest possible 
moment. A very large number of regular steel barrels 
w^as available for modification in the fall of 1944. The 
process of inserting stellite liners in regular steel bar¬ 
rels that had been developed by the Crane Company 80 
offered the best opportunity for immediate utilization 
of these barrels. This process is illustrated in Figure 9. 
Much of the success achieved was due to the selective 
fitting of ground liners with recessed barrels, for 
which careful gauging, such as shown in Figure 10, 
was required. 

The facilities used at the Crane Company (under 
Contract OEMsr-629) for making the 212 barrels for 
acceptance tests were expanded and an experimental 
production line was set up under a separate OSRD 
contract (OEMsr-1414) to develop efficient means of 
manufacture and prepare and deliver 2,000 stellite- 
lined barrels for combat tests. This work was com¬ 
pleted in slightly over 2 months on January 13, 1945, 
and these facilities were then taken over by the Ord¬ 
nance Department under an Army production con¬ 
tract. The 2,000 barrels were widely distributed to 
different theaters of operation, especially in the Paci¬ 
fic areas, and tested under combat conditions. Infor¬ 
mal reports of these tests confirmed the great superi¬ 
ority of this barrel over the standard steel barrel. Some 
of the barrels were tested by the Army Air Forces. 299 


h The standard steel barrel. 

* At this same meeting of the Ordnance Committee, the 
barrel (D-716948) having choked-muzzle chromium plate over 
a nitrided steel bore (developed by Division 1, as described in 
Chapter 23) was approved for manufacture as an alternate 
standard to stellite-lined barrels. 


CONFIDENTIAL 




452 


STELLITE-LINED MACHINE GUN BARRELS 



STEEL STELLITE 

RETAINING NUT LINED 
INSERTED MG BARREL 





Figure 9. The process of making a stellite-lined caliber .50 machine gun barrel from a standard steel barrel. 


Very early in the stellite liner development it was 
recognized by Division 1 that a method of manu¬ 
facture of stellite-lined barrels starting with barrel 
forgings was needed. Hence, negotiations with sev- 



Figure 10. Gauging of the diameters of rifled stellite 
liners which have been finish-ground to fit the tube re¬ 
cesses in the caliber .50 machine gun barrels. (Figure 
11 of NDRC Report A-447.) 


eral machine gun barrel manufacturers for develop¬ 
ment of a suitable means of doing so were initiated 
soon after the first successful accuracy and velocity 
firings were made. Before there was time for a con¬ 
tract to be negotiated, the Ordnance Department had 
decided to undertake this phase of the development. 
The gun barrel manufacturers, however, were reluc¬ 
tant to try such a radical departure from current 
manufacturing methods and were dubious concerning 
the practicability of machining operations on stellite. 

Accordingly, the Ordnance Department arranged 
a conference on October 10,1944, at the Philadelphia 
Office of Division 1 which was attended by represent¬ 
atives of all the prospective manufacturers of stellite- 
lined barrels. Representatives of the Division 1 staff 
and the Crane Company described the development 
of stellite liners and explained the procedures in use 
at Crane Company for pilot plant production. Fur¬ 
ther assistance was rendered to the Ordnance De¬ 
partment by having Crane Company supply detailed 
drawings of the special tools and operations to all 
prospective Ordnance contractors for stellite-lined 
barrels and by having representatives of prospective 
contractors visit and inspect the pilot plant produc¬ 
tion at Crane Company. 

The development of the procedures in use at Crane 
Company and a statistical analysis of the pilot plant 
production record on the 2,000 stellite-lined barrels 
are described in detail in a NDRC Report. 119 This 


CONFIDENTIAL 















































STELLITE APPLIED TO CALIBER .50 HEAVY BARREL 


453 


pilot plant production determined the best tools, 
machines, and sequence of operations for the recess¬ 
ing of steel barrels, boring and rifling of the liners, 
insertion of liners and alignment of rifling of the liner 
and rifled steel bore, preparation and chambering of 
the steel retaining nut, maintenance of tolerances and 
development of suitable inspection methods at all 
stages of the process and for the finished barrels. 
This demonstration at the Crane Company pilot 
plant of the practicability of manufacture of these 
barrels was an important contribution to the success 
of the stellite-lined barrel development. 

22 3 6 Studies to Facilitate 

Manufacturing Methods 

In order to facilitate the manufacture of stellite- 
lined barrels so that significant numbers of these 
barrels would get into immediate combat use, and 
also to insure the best utilization of Stellite No. 21 
as a bore-surface material, the research and develop¬ 
ment program of Division 1 included studies of the 
current manufacturing method and alternate meth¬ 
ods of applying this alloy as a liner or lining as well 
as the effects of various factors, such as variations in 
chemical composition, casting methods, heat treat¬ 
ment, mechanical-working, liner length, and thick¬ 
ness of liner wall, on the overall performance of lined 
barrels. 

In a few cases (only three were reported) the tack 
weld, used to prevent the steel retaining nut behind 
the stellite liner from unscrewing because of the vi¬ 
bration during firing, was brittle and the weld broke. 
In order to insure the reliability of this weld, studies 
of the welding process were made 80 which resulted in 
procedures that gave completely satisfactory welds. 
Studies were made of the effects of grinding versus 
reaming on the surface finish of the chamber, 80 of 
stellite-steel composite liners 89 (to eliminate the 
need for a retaining nut) and of the pressure bonding 
of stellite to steel. 89 

The methods for applying Stellite No. 21 to gun 
barrels by various infusion, incasting, and torch deposi¬ 
tion methods have already been described (Section 
19.4.3) and alternate methods of liner insertion in¬ 
cluding swaging into place, brazing, thermit (Al-Fin) 
bonding, and the Kelsey-Hayes method are described 
later (Section 26.2) when liner and other design fea¬ 
tures of gun tubes are discussed. The effects of vari¬ 
ous metallurgical factors on performance have already 
been discussed in Section 19.4.5. 


22 3 7 Liner Salvage 

Because of the possibility of a critical shortage of 
cobalt, the Ordnance Department informally re¬ 
quested Division 1 to investigate methods for the 
salvage of stellite liners from worn barrels. Crane 
Company investigated several methods 80 and found 
that (1) liners may be readily salvaged by the use of 
an oxy-acetylene cutting torch to cut the steel sur¬ 
rounding the liner into two sections (properly regu¬ 
lated depth of cut does not damage the liner); (2) 
pressing-out the liner by means of a hydraulic press 
at room temperature is unsatisfactory because of the 
excessive loads required; and (3) heating the outside 
of the barrel inductively and passing cold water 
through the bore enables most liners to be removed 
manually by means of a special extractor. The last 
method requires further study. 

22 4 APPLICATION OF STELLITE LINER 

TO THE CALIBER .50 HEAVY BARREL 

22.4.1 p re p ara Hon of Barrels and Test Results 

Although the first firing tests of Stellite No. 21 as a 
liner material were conducted in a caliber .50 heavy 
machine gun barrel (see Section 22.2.2) major atten¬ 
tion was given by Division 1 to the application of this 
liner material to improve the performance of the 
caliber .50 aircraft barrel because of the more urgent 
need for immediate improvement of this light (10 lb) 
barrel. The high level of performance of stellite-lined 
aircraft barrels described in Sections 22.3.3 and 22.2.2 
made it desirable to ascertain the effects of such a 
liner on the life and performance of the heavy (28 lb) 
barrel used with the ground machine gun. 

The Ordnance Department received a request from 
Headquarters, Army Ground Forces, submitted 
through Headquarters, Army Service Forces, for eight 
stellite-lined heavy machine gun barrels, four of them 
for test by the Antiaircraft Artillery Board and four 
by the Infantry Board. A total of 15 barrels was 
supplied by Division 1 for these tests and additional 
tests by the Ordnance Department, Army Service 
Forces. The results 295 are summarized in Table 1. 
They showed that inserting a stellite liner in the cal¬ 
iber .50 heavy barrel provides a considerable increase 
in accuracy life and for all practical purposes under 
normal firing conditions eliminates velocity drop as 
a consideration in the life of the heavy machine gun 
barrel. 


CONFIDENTIAL 





454 


STELLITE-LINED MACHINE GUN BARRELS 


Table 1. Army erosion tests on caliber .50 heavy ma¬ 
chine gun barrels. R: regular steel barrel; L: regular 
steel barrel containing a 9-in. liner of investment-cast 
Stellite No. 21. 295 


Schedule 

Burst Cooling* 

(rounds) (min) 

Bar¬ 

rel 

Life 

(rounds) 

Reason for 
end of life 

Vel. 

drop 

(fps) 

100 

2 

R 

885 

Tipping 

155 



L 

1191 

Tipping 

26 

5 X 100t 

1;5 

R 

1000 

Tipping 

183 



L 

2000 

Tipping 

1 

200 

1 or 2 

R 

1929 

Keyholing 

142 



L 

10554 

Keyholing 

124 

200 

Complete 

R 

3812 

Vel. drop and 

199 





dispersion 




L 

10225 

Dispersion 

2 


* Cooling in still air between bursts. 

t Five 100-round bursts with 1-min cooling between bursts and complete 
cooling after every 500 rounds. 


22,4,2 Adoption for Service Use 

As a result of these favorable firing results on 
barrels prepared by Division 1, the stellite-lined cal¬ 
iber .50 heavy machine gun barrel was adopted for 
Service use. At a meeting of the Ordnance Commit- 
t ee 29?s298 on May 31, 1945, recommendations were 
approved that “the barrel assembly (D-7161814) hav¬ 
ing a stellite liner be approved for immediate manu¬ 
facture as the preferred design for use in Gun, Ma¬ 
chine, Browning, Caliber .50, M2, Heavy Barrel, and 
that these barrels be marked in a distinctive manner,” 
that “the Stellite lined barrel (D-7161814) be placed 
in production as expeditiously as possible and super¬ 
sede the Barrel (D 28253-A) presently in manufac¬ 
ture,” and that “The Stellite lined barrels (D-7161814) 
be furnished one for each active caliber .50 M2 HB 
Gun, plus sufficient spare barrels to cover replace¬ 
ments.” 

Shortly after its adoption, contracts were award¬ 
ed by the Ordnance Department to Crane Company 
and several other barrel manufacturers for the mod¬ 
ification of regular steel heavy barrels by the inser¬ 
tion of 9-in. liners of investment-cast Stellite No. 21. 

22 4 3 Studies to 

Obtain Additional Performance 

Good evidence is presented later in this report (see 
Chapter 24) that the performance of aircraft machine 
gun barrels can be greatly enhanced by combinations 
of two or more of the following features (1) Stellite 
No. 21 breech liner, (2) choked-muzzle chromium 
plating of the steel bore surface ahead of the liner, 


(3) changes in the weight and external contour 
(distribution of weight) of the barrel, (4) use of an 
improved barrel steel with better high-temperature 
properties than regular WD 4150 machine gun steel, 
and (5) the use of internal or external barrel cooling 
methods. 

Studies were initiated by Division 1 at Crane Com¬ 
pany to evaluate the effects of such combinations on 
the performance of the caliber .50 heavy machine gun 
barrel. Thirty regular caliber .50 heavy (28 lb) barrels 
were modified by the insertion of regular 9-in. liners 
of investment-cast Stellite No. 21 and 12 additional 
barrels were prepared with both the liner and choked- 
muzzle chromium plating ahead of the liner. These 
barrels were to be used for a study of the effects of 
plate plus liner with and without variations in the 
weight and contour. No firing tests were made on 
these barrels before the termination of the experi¬ 
mental work of Division 1. Crane Company contin¬ 
ued these studies under an Ordnance Department 
research and development contract. 

22 5 APPLICATION OF A STELLITE LINER 
TO CALIBER .30 BARRELS 

22,51 Development of the 

Design and Manufacture 

As the result of an urgent request from the Army 
Ground Forces to the Ordnance Department, Divi¬ 
sion 1 was asked to apply a liner of Stellite No. 21 to 
caliber .30 machine gun barrels. Crane Company was 
asked to prepare a design for the application of a stel¬ 
lite liner to the caliber .30 aircraft machine gun bar¬ 
rel, M2, and to prepare a few barrels to test the design. 
This design was essentially similar to that employed 
for caliber .50 barrels (see Figure 1). In this case, how¬ 
ever, the flanged investment-cast liner was lighter 
and was 6 inches in overall length. Two barrels tested 
at the Crane firing range showed that the design was 
satisfactory and that these stellite-lined barrels were 
far superior in accuracy and velocity life to standard 
unlined barrels tested for comparison. Crane Com¬ 
pany delivered ten additional barrels for Ordnance 
Department tests. 

It was desirable to supply the Ground Forces for 
immediate use improved barrels with as low a barrel 
weight as practical. Test results were needed on three 
caliber .30 barrels—those used in aircraft M2, the 
M1919A6 and the M1917A1 guns, with barrel 
weights of approximately 3.8, 4.5, and 3 lb, respec- 


CONFIDENTIAL 







STELLITE APPLIED TO CALIBER .30 BARRELS 


455 


tively. The last barrel is normally used in a water- 
cooled gun but was to be tested for use without water¬ 
cooling in the M1919A6 or M1919A4 gun by means 
of an adapter. To expedite early Service application 
if results of test were satisfactory, Division 1 nego¬ 
tiated contracts with two caliber .30 barrel manufac¬ 
turers (Johnson Automatics Company and Reming¬ 
ton Arms Company) to apply a stellite liner “by a 
practical method” and deliver barrels for test by the 
Ordnance Department. 

Each of the contractors inserted stellite liners sup¬ 
plied by the Haynes Stellite Company in 50 aircraft 
M2 barrels, 50 M1919A6 barrels, and 25 M1917A1 
barrels, and delivered the lined barrels for Ordnance 
Department tests. Some preliminary testing of bar¬ 
rels was done by each contractor to prove up various 
new design features and indicate the order of magni¬ 
tude of the superiority of lined barrels over standard 
steel barrels. The designs and methods used by these 
contractors are described later in this report (Sec¬ 
tions 26.2.3 and 26.2.4) where liner and design fea¬ 
tures of gun tubes are discussed more fully. 

In addition to the efforts just described, Division 1 
cooperated with the Small Arms Division, Research 
and Development Service of the Ordnance Depart¬ 
ment in the following ways. Crane Company was 
asked to insert stellite liners and deliver three barrels 
for a Garand-type caliber .30 automatic rifle. No in¬ 
formation has been received by Division 1 on the re¬ 
sults of tests on these barrels. The Johnson Auto¬ 
matics Company was asked to prepare 12 lined bar¬ 
rels for the Johnson light machine gun Ml945, 4 lined 
barrels M1919A6 modified for interchangeable adap¬ 
tion with the M1919A4 gun using the T27 flash hider 
and muzzle cap, and 6 lined M1919A6 barrels modi¬ 
fied by maintaining the maximum dimension at the 
breech forward to a point 2 in. beyond the liner. 

21,5 2 Test Results 

The results 127 of firing tests conducted for Division 
1 at Johnson Automatics Company on stellite-lined 
caliber .30 machine gun barrels may be summarized 
as follows: 

M2 Aircraft: In a continuous burst of 500 rounds 
at 1,200 rounds per minute, the lined barrel was ac¬ 
curate throughout the 500 rounds, whereas the stand¬ 
ard steel barrel showed complete keyholing at 360 
rounds. In 250-round bursts at 1,200 rpm with 15 min 
air cooling, the lined barrel lasted 5,000 rounds, the 
standard barrel 1,200 rounds. In 300-round bursts at 


800 rpm with 2 min air cooling, the lined barrels 
lasted 950 rounds, the standard barrels 450 rounds. 

M1919A6: In 250-round bursts at 550 to 600 rpm 
with 2 min air cooling, the lined barrels lasted 750 to 
800 rounds, the standard barrels 350 rounds. In 100- 
round bursts with 30 sec air cooling, the lined barrel 
lasted 1,800 rounds, the standard barrel 950 rounds. 

M1917A1: When a 3-lb lined barrel was fired in 
the air-cooled M1919A4 gun with an adapter, on a 
schedule of 250-round bursts with 2 min cooling in 
comparison with the 7.5-lb heavy M1919A4 standard 
steel barrel, each barrel had a life of 600 rounds, but 
the lined barrel was badly bulged in this test. 

M1919A6; Modified Contour: When two lined bar¬ 
rels, further modified by maintaining the largest diam¬ 
eter of the breech to a point 2 in. beyond the liner (ad¬ 
ditional weight 0.5 lb) were fired in 250-round bursts 
with 2 min air cooling, they lasted over 1,150 rounds 
each as compared to 750 to 800 rounds with lined bar¬ 
rels of standard external contour. 

Johnson M191+5: In continuous fire at a delivered 
rate of 200 rpm (cyclic rate per 20-shot bursts 750 to 
800 rpm) the lined barrel lasted over 700 rounds, the 
standard steel barrel 350 rounds. 

The results 135 of firing tests conducted for Division 
1 at Remington Arms Company may be summarized 
as follows: 

M2 Aircraft: In 300-round bursts with 2 min cool¬ 
ing, 13 lined barrels showed an average accuracy life 
of 1,243 rounds. The poorest barrel had a life of 663 
rounds and the best barrels had not lost accuracy 
when firing was discontinued after 2,000 rounds. In 
comparison, 11 standard steel barrels averaged 486 
rounds accuracy life on this schedule (poorest 400, 
best 726 rounds). 

M1917A1: Seven lined barrels when fired in the 
M1917A1 water-cooled gun with no water in the 
jacket on the schedule of 300-round bursts with 2 min 
cooling gave a life of between 300 and 600 rounds 
(average 483 rounds). These barrels all failed due to 
bulging of the barrel at the muzzle end of the liner. 
The standard steel barrels failed for accuracy be¬ 
tween 230 and 440 rounds (average 312). Improved 
performance (average accuracy life 635 rounds) and 
no bulging occurred in a test of four lined barrels pre¬ 
pared with a straight cylindrical stellite liner with 
0.430-in. outside diameter. 

The results of Ordnance Department tests 297 - 320 on 
the stellite-lined M1919A6 barrels are given here as 
Table 2. These data indicate that the stellite-lined 
caliber .30 barrel has an accuracy life two to three 


CONFIDENTIAL 



456 


STELLITE-LINED MACHINE GUN BARRELS 


Table 2. Summary of result of Ordnance Department 
firing tests of stellite-lined caliber .30 machine gun 
barrels, M1919A6, supplied by Division 1, NDRC, 
compared with unlined barrels. 


Firing cycle 
Burst Cooling* 
(rounds) (sec) 

Kind 

of 

barrel 

Number Accuracy- 
barrels lifef 

fired (rounds) 

Velocity 

change 

(fps) 

20 

10 

Lined 

2 

1700 

+31 



Unlined 

3 

613 

-55 

200 

120 

Lined 

2 

1080 




Unlined 

3 

583 


300 

120 

Lined 

5 

759 

+46 



Unlined 

3 

448 

-80 

400 

120 

Lined 

2 

720 




Unlined 

2 

380 



* In still air. 

f Rounds to serious key holing. 


times as great as that of the regular steel barrel. 
Furthermore, the lined barrels suffered no permanent 
loss in velocity but instead an appreciable increase 
even though the firing schedules used were extremely 
severe. In all tests the functioning of the guns was 
normal. 

Just as in caliber .50 aircraft machine gun barrels 
(see Section 22.3.3) caliber .30 barrels with liners 
show a greater improvement over standard steel bar¬ 
rels on severe firing schedules than on mild schedules, 
although even on mild schedules the improvement is 
substantial. 

22,5,3 Adoption of Lined M1919A6 Barrel 
for Service Use 

As a result of the very favorable test results on the 
stellite-lined caliber .30 M1919A6 barrels delivered 
by Division 1, this barrel was adopted for Service use. 
At a meeting of the Ordnance Committee 297 ’ 298 on July 
26, 1945, recommendations were made and approved 
that “The Barrel Assembly (Drawing No. D-7162295) 
having a Stellite liner be approved for immediate 
manufacture as the preferred design for use in Gun, 
Machine, Caliber .30, Browning, M1919A6, and that 
these barrels be marked in a distinctive manner, and 
issued with highest priority to units now in combat,” 
that “The Stellite-lined Barrel (D-7162295) be placed 
in production as expeditiously as possible and super¬ 
sede the Barrel* (D-54559) presently in manufacture,” 
and that “The Stellite-lined barrels (D-7162295) be 
furnished one for each active M1919A6 caliber .30 
Machine Gun, plus sufficient spare barrels to cover 


j Standard steel barrel for caliber .30 gun M1919A6. 


replacements.” Plans for large-scale manufacture 
were initiated but were not carried out owing to cut¬ 
backs in the procurement program following V-J Day. 

Although the performance of the caliber .30 ma¬ 
chine gun barrel, aircraft, M2, was substantially in¬ 
creased by the use of a stellite liner, this weapon was 
obsolete for aircraft armament and had been sup¬ 
planted by the caliber .50 aircraft machine gun with 
its superior fire power. The immediate use of caliber 
.30 aircraft barrels by Ground Forces would require a 
modification of mounts and would necessitate a large 
procurement and supply program for guns, mounts 
and other accessories and delay the use of improved 
barrels by Ground Forces in combat. The tests on the 
very light (3 lb) caliber .30 M1917A1 barrel with a 
liner had shown failure by bulging due to overheating, 
aggravated by the thinness of the steel barrel wall at 
the forward end of the liner. 

Studies of the application of the stellite liner to 
automatic rifles and other caliber .30 barrels was con¬ 
tinued by the Ordnance Department at Crane Com¬ 
pany after Division l’s contract was terminated. 

22 6 APPLICATION OF A STELLITE LINER 
TO THE CALIBER .60 BARREL 

22,6,1 Design and Performance 

During World War II, the Ordnance Department 
was developing a new aircraft machine gun with a 
caliber .60 bore, a high cyclic rate of fire, and a very 
high projectile velocity (slightly above 3,500 fps) to 
deliver more fire power at higher velocities than in 
any previous aircraft weapons. In early tests the 
Ordnance Department found that steel barrels would 
survive only a few hundred rounds owing to severe 
breech erosion and very rapid loss in velocity life. 
Since a considerable improvement in barrel life was 
imperative in order to make the weapon practical, 
the Ordnance Department requested Division 1 to 
apply a stellite liner to this barrel. 

Crane Company suggested a design for the inser¬ 
tion of a 12-in. length flanged breech liner of Stellite 
No. 21 on a shrink-fit, a design essentially similar to 
that successfully used in the caliber .50 aircraft ma¬ 
chine gun barrel (see Section 22.2) and prepared 
some lined barrels. Tests at Aberdeen Proving 
Ground of the first two lined barrels showed the 
outstanding performance of these barrels. On the 
same test schedule as previously used on regular 
steel barrels for this gun, after 500 rounds (standard 


CONFIDENTIAL 








STELLITE APPLIED TO CALIBER .60 BARREL 


457 


ammunition with IMR powder) a velocity drop of 
only 30 fps occurred with the lined barrels instead 
of one of 400 fps which had been encountered with 
steel barrels. Also on this schedule the lined barrels 
showed a 200-fps velocity drop only after about 
2,000 rounds. No inaccuracy was encountered. Thus 
the use of a stellite liner in this barrel increased the 
life about sixfold on a mild schedule and made pos¬ 
sible the firing of 50-round bursts, which were en¬ 
tirely too severe for a steel barrel. Crane Company 
prepared and delivered a total of 70 stellite-lined 
barrels for this gun, some of them with fluted 
chambers as described below. The stellite-lined barrel 
completely supplanted steel barrels for this weapon. 

In addition to the regular caliber .60 barrels, Divi¬ 
sion 1 cooperated with the Ordnance Department by 
having Crane Company insert stellite liners into six 
Mann barrels (three of which were provided with cut 
flutes in the chambers) for tests at Frankford Arsenal 
to determine the effect of erosion ahead of the liner on 
the performance of explosive ammunition and on 
liners inserted in two barrels of an experimental mod¬ 
el, caliber .60, T2E4. Difficulties with ammunition 
prevented the satisfactory testing of these last two 
barrels. 

In view of the success obtained with caliber .50 air¬ 
craft barrels by a combination of stellite liner with 
chromium-plating ahead of the liner (described in 
Chapter 24), two caliber .60 barrels were prepared 
with both a stellite liner and chromium plate ahead of 
the liner, and delivered to the Ordnance Department. 

22 6 2 Chamber Fluting 

Difficulties with the gun mechanism and ammuni¬ 
tion prevented early Service use of caliber .60 barrels 
with a liner. During the firing of caliber .60 barrels k 
considerable difficulty with extraction of cartridge 
cases was encountered. The Ordnance Department 
found it necessary to provide longitudinal flutes in 
the forward half of the chamber to admit gas behind 
the cartridge case to aid in releasing it after firing. 


k All firing tests on caliber .60 barrels were conducted by the 
Ordnance Department and none were made by Division 1, 
NDRC, who supplied lined barrels for test. 


Fifty-three of the total 70 lined barrels delivered for 
test by Crane Company just described were fitted 
with steel retaining nuts made from the fluted cham¬ 
bers of regular steel barrels. The chamber section of 
the liners in these barrels were fluted by a punch-type 
of tool, and the retainers of all but the first two bar¬ 
rels were pinned to prevent rotation. 

During the testing of lined barrels with fluted 
chambers, it became evident that the number and 
depth of flutes then being used were not necessarily 
the most suitable for the purpose and that there was 
an urgent need for further tests to determine optimum 
flute dimensions. Crane Company designed and con¬ 
structed a special hand-operated fluting tool with 
which the flute dimensions could be varied at will. In 
tests at Aberdeen Proving Ground of flutes in ten un¬ 
lined barrels prepared at Crane Company by the use 
of this tool, only the barrels with 16 flutes 0.040-in. 
deep could be extracted without failure. Subsequently 
lined barrels with similar flutes were sent to the 
Ordnance Department for test. 

liner with Integral Chamber 

Machining difficulties encountered in fluting of 
composite steel-stellite chambers and difficulties with 
the erosion that was found to occur in fluted steel 
chambers led to the design at Crane Company of an 
integral liner and chamber made of investment-cast 
Stellite No. 21. In this design the liner flange or 
shoulder is extended to the breech end of the barrel 
and no steel retainer is used. This design should have 
the advantages of easier fluting, no joint in the cham¬ 
ber, erosion-resistant flutes, simplified machining- 
through elimination of the retainer, possibility of 
precision casting the flutes into the chamber, and 
the possibility of eliminating the flutes entirely 
because of the lower coefficient of friction of stellite. 
Investment-cast liners were prepared by the Haynes 
Stellite Company under Contract OEMsr-1330 and 
delivered to Crane Company for insertion into caliber 
.60 barrels under Contract OEMsr-629. The prepara¬ 
tion of the lined barrels was started but not com¬ 
pleted before termination of experimental work by 
Division 1. Some of these barrels should be com¬ 
pleted and tested. 


CONFIDENTIAL 




Chapter 23 

NITRIDED AND CHROMIUM-PLATED MACHINE GUN BARRELS 

By E. F. Osborn h 


« 1 DEVELOPMENT OF THE NITRIDED, 
CHROMIUM-PLATED CALIBER .50 AIR¬ 
CRAFT BARREL WITH CHOKED MUZZLE 

2,11 Introduction 

D ivision 1 pursued two parallel problems for the 
improvement of machine gun barrels by the ap¬ 
plication of erosion-resistant materials. The develop¬ 
ment of barrels with a stellite liner has just been 
described in Chapter 22. Similar success with caliber 
.50 aircraft barrels was achieved by use of chromi¬ 
um electroplate on a hardened steel bore. The de¬ 
velopment by the Geophysical Laboratory and the 
National Bureau of Standards of the nitrided, chromi¬ 
um-plated caliber .50 aircraft barrel with a choked 
muzzle is described in this chapter. Both barrels were 
adopted 294 296 at the same time to supersede the reg¬ 
ular steel barrel for Service use, the stellite-lined as 
the preferred barrel and the nitrided and chromium- 
plated barrel as an alternate. The need of our Air 
Forces and those of our allies for an improved caliber 
.50 aircraft barrel which would fire accurately 
throughout long bursts of rapid fire is discussed in 
Chapter 22. 

As soon as Division 1 became fully aware of this 
need, which was early in 1944, the development of 
the two types of improved barrels was initiated and 
proceeded rapidly as a result of its previous research 
on the fundamental aspects of erosion of steel and 
chromium-plated steel guns and on means of mitigat¬ 
ing erosion by the use of erosion-resistant materials. 
Pilot plant production of the improved barrels was 
undertaken by the W. B. Jarvis Co., c as described in 
Chapter 25. 

Experiments with plated steel liners in the caliber 
.50 heavy barrel d showed that if chromium plate, 

* This chapter has been prepared by C. A. Marsh by ar¬ 
rangement of appropriate sections from NDRC Report 
A-409 81 by E. F. Osborn, following an outline suggested by 
J. F. Schairer. 

b Petrologist, Geophysical Laboratory, Carnegie Institution 
of Washington. (Present address: Department of Earth Sci¬ 
ences, Pennsylvania State College, State College, Pa.) 

c Now the Jarvis Division of the Doehler-Jarvis Corporation. 
d “Heavy” and “aircraft” caliber .50 barrels are defined in 
Section 11.2.2. 


which has excellent erosion resistance, was to remain 
near the origin of rifling long enough to improve the 
velocity-life of this barrel, the metal of the lands 
under the chromium had to be more resistant to 
swaging (Section 10.5.3.) than WD 4150 steel. 80,85 
Improvement in this respect was obtained by harden¬ 
ing this same steel by nitriding. One big advantage of 
using WD 4150 steel in the development of improved 
barrels was that a large stockpile of barrels made of 
this steel was on hand. By nitriding and properly 
chromium-plating them, an improved barrel could 
therefore be obtained through the use of existing 
stocks. 

Experiments with plated steel liners in the heavy 
barrel had shown that the softer, LC (low contrac¬ 
tion) chromium resisted removal better than the 
hard, HC (high contraction) chromium, e as men¬ 
tioned in Section 16.3.8. Adhesion of this soft plate to 
steel, however, when applied to the full length of the 
aircraft barrel, was never consistently good, and it 
was removed from the bore as rapidly, or often more 
rapidly, than the hard chromium, which was there¬ 
fore standardized for the aircraft barrel. 

Inasmuch as neither nitriding nor chromium-plat¬ 
ing involved critical processes or materials, it was 
decided to treat the full length of the aircraft barrel 
rather than use a liner. The principal steps in the 
process of transforming a standard, steel caliber .50 
machine gun barrel into a nitrided and chromium- 
plated barrel with a choked muzzle are illustrated in 
Figure 1. 

It was found that the main improvement of this 
aircraft barrel was in accuracy-life, as is brought out 
in Section 23.1.4. The accuracy-life was improved 
largely because the muzzle section of the bore was 
choked and because this choke remained on firing 
owing to the lack of erosion of the chromium surface. 
In addition, the velocity-life was improved when 
severe firing schedules were used because chromi¬ 
um stayed on the bore surface at the origin of 
rifling long enough to serve as a protection to the 
steel, and because nitriding of the steel minimized 
swaging of the lands. 

e For the properties of the different types of chromium plate, 
see Section 20.2.2. 


458 


CONFIDENTIAL 




THE NITRIDED, CHROMIUM-PLATED AIRCRAFT BARREL 


459 


NITRIDING 

TO HARDEN 
THE BORE 


ELECTROPOLISHING 
THE BORE 


CHROMIUM PLATING 

THE/ BORE 

( Tapered thickness to^ 
give a choked muzzle) 




Figure 1 . The process of making a nitrided and chromium-plated caliber .50 machine gun barrel. 


2,1,1 Development of Specifications for 
Nitrided and Chromium-Plated Barrels 

Dimensions 

Interior dimensions of the finished nitrided, chro¬ 
mium-plated barrels were continually changed as re¬ 
search continued, the changes being made either to 
improve the performance of the barrel or to simplify 
production where the change did not adversely affect 
performance. The trend of the changes was to de¬ 
crease the bore diameter at the muzzle and to in¬ 
crease it at the breech, and to increase the tolerance 
at the muzzle end. The development of the final spec¬ 
ifications for dimensions is given in detail in Section 
25.2.2. 

Nitriding 

Nitriding of the barrels was done by both gas 
(ammonia) and liquid-immersion methods, the for¬ 
mer being used exclusively in production during the 
war. Research was done on various times and tem¬ 
peratures of nitriding. 109 - 110 Nitriding, by a cycle 
common at the Link Belt Company, Philadelphia, 
was adopted as being satisfactory. The barrels, stand¬ 
ing vertically, were held at 950 F for 38 hr. All sur¬ 


faces of the barrel were nitrided. The hard case on 
the outside was actually a slight advantage, adding a 
little strength to the barrel. The incidental nitriding 
of the breech threads or the chamber was found to be 
no disadvantage. 

Before nitriding, the barrel was cleaned free of 
grease and decoppered if it had been proof-fired. 
Later, the barrels to be chromium plated at the pilot 
plant described in Chapter 25 were also tested for 
hardness and straightness prior to nitriding; if the 
hardness on the outside was below 28 Rockwell C, 
the barrel was not used. In the initial experimental 
plating a stress-relieving procedure preceding nitrid¬ 
ing was used, but this was abandoned when it was 
found to be of no significant advantage. 

Inasmuch as nitriding processes differ somewhat 
from plant to plant, and the initial hardness and com¬ 
position of the steel affects the hardness and depth of 
the case, it was desirable to set up a specification 
regarding these two characteristics of a case. 

In general, the higher the initial, or core, hardness 
of the steel, the harder is the case, both near the sur¬ 
face and at depth. The carbon content also affects the 
hardness. For initial hardness of 29 and 36 Rockwell 
C, the hardness of the case at 0.001 in. below the sur¬ 
face is of the order of 500 and 650 microvickers, 


CONFIDENTIAL 






























































460 


NITRIDED AND CHROMIUM-PLATED MACHINE GUN BARRELS 


respectively, for WD 4150 steel nitrided 38 hr at 
950 F. 

Measurements to determine hardness and depth of 
case were made using the Eberbach microvickers 
hardness tester. 110 Using a polished cross section of 
the barrel, a hardness profile of the case was obtained. 
A typical hardness profile after nitriding 38 hr in 
ammonia is shown in Figure 2. The tentative specifi¬ 
cation set up was that the hardness should be a 
minimum of 500 VPN at a depth of 0.001 to 0.002 in. 
below the surface, and a minimum of 400 VPN at a 
depth of 0.010 in. 



DISTANCE FROM BORE IN INCHES 

Figure 2. Hardness profile of nitrided case in a caliber 
.50 barrel before firing. Barrel No. L2-J45, nitrided 38 
hours in ammonia by Link Belt Company. (This figure 
had accompanied the manuscript of NDRC Report 
A-409 but was not reproduced with that report.) 

Electropolishing and Chromium-Plating 

The procedure for electropolishing and chromium¬ 
plating caliber .50 aircraft barrels to give the bore 
dimensions and thickness and quality of chromium 
which were desired was worked out largely by the 
Electrochemical Section at the National Bureau of 
Standards 84 and is described in Chapter 20. Improve¬ 
ments were made by the Doehler-Jarvis Corporation 
during the pilot-plant stage in which an extensive 
research program was carried out to try to evaluate 
the importance of various steps in the plating proced¬ 
ure. Before large-scale production got under way 
with the attendant, necessary freezing of the pro¬ 
cedure, it was desirable to eliminate all steps found to 
be unnecessary, change any of the steps in the direc¬ 


tion of improving the adhesion of the chromium, and 
evaluate the criticalness of the various specifications. 
On the basis of this research, described in Section 
25.2, the final specifications and procedure were 
established. 

231,3 Firing Tests 

Firing Schedules 1 

Firing schedules are always a compromise if one 
wants to obtain all possible information on a barrel. 
If, for example, one fires a continuous burst to key- 
holing, 8 he learns only at what stage bullets begin to 
keyhole. Inasmuch as the test is not interrupted until 
the barrel is essentially worn out, there is no chance 
for observing progressive erosion in the barrel and for 
obtaining other valuable information. 

In general, a barrel should be fired in a manner sim¬ 
ilar to its use in combat. But this is not alwaj^s prac¬ 
ticable or even desirable for several reasons: (1) It is 
very difficult to decide how a gun is typically fired in 
combat; the same gun may be fired very differently, 
for example, on an air-to-ground mission than on an 
air-to-air mission, or on the same type of mission in 
different theaters; (2) a testing schedule has to be 
one that can be executed in a short time; (3) a firing 
schedule has to be arranged so that observations can 
be made during the progress of the test, and many of 
these observations can be made only on the cold 
barrel. 

Consequently, there are valid objections raised to 
every firing schedule. One has first to decide what it 
is most necessary to test in a barrel, then by trial find 
a schedule which will do this reasonably well and at 
the same time meet the requirements that it not be 
too far removed from battle practice and that it per¬ 
mit obtaining as much desirable additional informa¬ 
tion on the performance of the barrel as possible. 

In all, 13 different firing schedules were used by the 
Geophysical Laboratory as the need arose at various 
times during its program for comparing caliber .50 
barrels, and a few other tests such as a 500-round 
burst were used for special barrels. Each of the firing 
schedules was given an identifying symbol, such as 
5 X 100 (2). The schedules will be briefly described 


f Some of these same firing schedules were used in the test¬ 
ing of stellite-lined caliber .50 barrels, with the results given 
in Section 22.3.3. 

g See footnote (e) in Chapter 22. 


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THE NITRIDED, CHROMIUM-PLATED AIRCRAFT BARREL 


461 


below, and then later they will be referred to only by 
symbol. 

GL-135 Schedule. This was the first schedule stand¬ 
ardized and it was used only for testing heavy bar¬ 
rels. This schedule was not intended necessarily to 
bear any relation to the manner in which the caliber 
.50 heavy barrel gun was fired in combat, but was 
carried out to obtain information on erosion which 
might be applicable to large antiaircraft guns. It was 
desired (1) to have as many of the rounds as feasible 
pass through the barrel while it was approximately at 
constant temperature; (2) to wear out a standard 
steel barrel within a reasonable time, the rejection 
limit being set at a 200-fps drop in velocity, and (3) 
to have the rounds fired in several similar groups of 
bursts so that four or five times during the test the 
barrel could be cooled for examination and for 
gauging. 

After several firing trials of barrels equipped with 
thermocouples for measuring the temperature during 
firing at a point Ms-in. from the bore surface at the 
origin of bore, as described in Section 5.3.6, it was de¬ 
cided to fire groups of 400 rounds to a total of 2,000 
rounds. Each group consisted of an initial 135-round 
burst followed by a series of shorter bursts, the length 


of burst and interval between bursts being estab¬ 
lished so that in the standard barrel the temperature 
at the point near the origin of bore would remain, 
after the initial burst, between 450 and 500 C: Other 
points in the barrel would not, of course, necessarily 
remain at a constant temperature. Temperature 
curves for various points in the barrel during this 
firing schedule are shown in Figure 3. 

5 X 50 {1) Schedule. A 250-round group was fired, 
consisting of five 50-round bursts, one burst being 
fired each minute. The barrel was cooled to ambient 
temperature and the schedule repeated. 

5 X 100 (1) Schedule. Same as above except that 
100-round bursts were used instead of 50-round 
bursts. 

5 X 120 ( 1) Schedule. Same as above except 120- 
round bursts were used. This schedule, used for heavy 
barrels, is comparable from the standpoint of barrel 
temperature to the 5 X 100 (2) schedule in an air¬ 
craft barrel. 

5 X 100 (2) Schedule. Groups of five 100-round 
bursts were fired with an interval of 2 min between 
the end of each burst and the beginning of the next, 
except that complete cooling to ambient temperature 
was allowed between groups. 



Figure 3. Temperature-time curves obtained during firing of GL135 schedule (400 rounds) in a caliber .50 heavy 
barrel. (Note: Number of rounds in a burst shown by scale in upper left of graph.) (Figure 5 in NDRC Report A-409.) 


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462 


NITRIDED AND CHROMIUM-PLATED MACHINE GUN BARRELS 


10 X 1+0 (1) Schedule. In this schedule, used com¬ 
monly by the British, ten 40-round bursts were fired 
at 1-min intervals. The barrel was cooled and the 
schedule repeated. 

20 X 25 ( 3 ) Schedule. Twenty 25-round bursts were 
fired at intervals of 3 min. The barrel was cooled to 
ambient temperature and the schedule repeated. 

1 X 300 Schedule. A single 300-round burst was 
fired, the barrel cooled to ambient temperature, and 
the burst repeated. 

1 X 150 Schedule. Same as above but with a 150- 
round burst. 

1 X 100 Schedule. Same as above, but with a 100- 
round burst. 

CGL-350 Schedule. This schedule consisted of an 
initial 350-round burst followed, after the barrel was 
cooled to ambient temperature, by a series of 5 X 100 
(2) groups. The “CGL” stands for Crane-Geophysical 
Laboratory. The schedule, the development of which 
is described in Section 24.1.3, was used almost ex¬ 
clusively by these two contractors of Division 1 dur¬ 
ing the last few months of the war in testing the 
“combination” caliber .50 aircraft barrels described 
in Chapter 24. 

O BURST AMMUNITION AP-M2 LOT TW 19031 
• BURST AMMUNITION AP-M2 LOT M 29558 


ALL BARRELS FIRED 5XIOO (2) SCHEDULE 



700 750 800 850 900 950 1000 

CYCLIC RATE IN ROUNDS PER MINUTE 


Figure 4. Change in muzzle velocity versus cyclic rate 
for nitrided, chromium-plated caliber .50 aircraft barrels 
with choked muzzles, fired with two different lots of 
ammunition. (Figure 19 in NDRC Report A-438.) 



Figure 5. Targets for nitrided, chromium-plated cal¬ 
iber .50 aircraft barrel No. 49, during the third erosion 
group. (This figure had accompanied the manuscript of 
NDRC Report A-409 but was not reproduced with that 
report.) 

C-l Scheduled The barrel was fired continuously 
until serious keyholing of the bullets occurred. 

B-l Schedule . h The barrel was fired in 100-round 
bursts separated by 2-min intervals until serious key- 
holing of the bullets occurred. 

Ammunition 

The type of caliber .50 ammunition used in testing 
depended on the object of the test. The following are 
the standard types that were used: ball-M2; ar¬ 
mor piercing, AP-M2; armor piercing incendiary, 
API-M8 andAPI-M8E1I; incendiary, I-Ml, I-M1E1, 

h The behavior of standard steel barrels when fired according 
to this schedule, which is an Army Ordnance Department 
schedule, is described in Section 22.3.3. 


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THE NITRIDED, CHROMIUM-PLATED AIRCRAFT BARREL 


463 



Figure 6. Targets for nitrided, chromium-plated cal¬ 
iber .50 aircraft barrel No. 49, during the eleventh and 
twelfth erosion groups. (This figure had accompanied 
the manuscript of NDRC Report A-409 but was not re¬ 
produced with that report.) 


and I-M23 (experimental I-T48) tracer, Tr-Ml, 
Tr-MlO, and Tr-MlOEl Rounds designated by the 
suffix “El” were loaded with double-base powder 
containing 20% nitroglycerin, whereas all the others 
were loaded with a single-base powder, namely 
IMR 5010, which is less erosive than the double¬ 
base powder. Several types of specially loaded am¬ 
munition were also used in some tests. 

The bullets in these rounds of ammunition have 
gilding metal jackets, annealed or unannealed. Their 
different types of construction affect their resistance 
to engraving, and thus had a perceptible effect on the 
rate of erosion of the barrels that were tested. The 
melting and softening of the jackets at the tempera¬ 
tures attained during firing long bursts may definitely 
limit machine gun performance. 

Whenever it is desired to make as accurate as pos¬ 


sible a comparison of different barrels, a single type of 
ammunition and if possible a single lot of that am¬ 
munition should be used. A “combat” load, where 
two or three types of ammunition are used, is not to 
be recommended. Either ball-M2 or AP-M2 ammuni¬ 
tion was used in most of the testing of chromium- 
plated barrels. Even with AP-M2 ammunition, where 
one might expect very little variation from lot to lot, 
considerable variation was found. 110 Figure 4 illus¬ 
trates the effect of two factors on velocity drop. The 
plot of cyclic rate against velocity change possibly 
shows but slight correlation between rate of fire and 
velocity change. There is, however, a definite cor¬ 
relation between lot of ammunition and velocity 
drop. 


Velocity Measurements 


Velocity measurements were made at the begin¬ 
ning of and at various stages during the firing tests by 
means of Aberdeen chronographs with screens at 25 
and 75 ft from the muzzle. Both hot and cold veloc¬ 
ities were determined. The former refer to measure¬ 
ments taken at the end of a burst before the barrel 
was allowed to cool, while the latter refer to measure¬ 
ments with the barrel at ambient temperature. The 
rounds used for determining the cold velocities were 
held at a constant temperature for a minimum of 16 
hr preceding the test in order to eliminate the effect 
of differences in the initial powder temperature on 



Figure 7. Targets for standard, caliber .50 aircraft 
barrel No. 58, during the first erosion group. (This fig¬ 
ure had accompanied the manuscript of NDRC Report 
A-409 but was not reproduced with that report.) 


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464 


NITRIDED AND CHROMIUM-PLATED MACHINE GUN BARRELS 



Figure 8. Targets for standard caliber. 50 aircraf t bar¬ 
rel No. 58, during the second erosion group. (This 
figure had accompanied the manuscript of NDRC Re¬ 
port A-409 but was not reproduced with that report.) 


Figure 9. Targets for standard caliber .50 aircraft 
barrel No. 237. Bullets began keyholing in the second 
100-round burst. (Figure 6 in NDRC Report A-409.) 


velocity of the bullet. They were fired after five 
tvarm-up rounds. 

Accuracy Measurements 

Targets. The two principal measurements were dis¬ 
persion, or size of target, and yaw of bullets, or angle 
between axis of bullet and direction of flight. The 
standard distance from muzzle to target was 1,000 in. 
A second target was sometimes placed at 2,000 in. 
For most firing tests, fixed targets were used. Hence, 
during the firing of schedules like 5 X 50 (1), 5 X 100 
(2), and B-l, a new target was set up for each burst. 
Figures 5 through 10 are photographs of examples of 
such targets. 

A moving screen was used whenever it was desired 
to know just what rounds in a burst “tipped” or 
“keyholed.” In the recording of data, a yaw of 20 de¬ 
grees, for which the length of hole is 1 in., was arbi¬ 
trarily set up as a critical measurement, and the tar¬ 
get holes were recorded as being made by bullets hav¬ 
ing a yaw of either less than or greater than 20 de¬ 
grees. “Tippers” may have a yaw angle less than this. 



Figure 10. Targets for nitrided, chromium-plated 
caliber .50 aircraft barrel No. JN3. (Figure 7 in NDRC 
Report A-409.) 


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THE NITRIDED, CHROMIUM-PLATED AIRCRAFT BARREL 


465 


Collection of Bullets. A study of engraving on col¬ 
lected bullets often gave information that was not 
obtainable from target data. The angle of engraving 
on a bullet gives an indication of the amount of spin 
and therefore of the stability of the bullet. A bullet 
having an angle of engraving of 6 degrees (full spin) 
will normally straighten out even though at 1,000 in. 
from the muzzle it has a slight yaw. Therefore when a 
collected bullet has an angle of engraving of 6 degrees, 
it can reasonably be assumed that it was stable, even 
if it was yawing at 1,000 in. Examples of collected 
bullets are shown in Figure 11. Only the first two did 
not keyhole. The third keyholed even though the 
angle of engraving was approximately 6 degrees. 

231,4 Performance of Nitrided and 
Chromium-Plated Barrels 

Hundreds of firing tests were made on the nitrided, 
chromium-plated barrels with choked muzzles. In 
very few of these tests, however, were the barrels 
fired to the end of accuracy-life and ordinarily were 
not even fired to end of velocity-life, for the firing was 
done largely to study the effect of some change in the 
method of producing the barrel on the adhesion of 
chromium, resistance to swaging of the lands, engrav¬ 
ing of the bullet, etc. However, enough barrels were 
fired to the end of life using various firing schedules 
that a good picture of the performance of these barrels 
was obtained. In addition to the barrels fired at the 
Geophysical Laboratory, some were fired at Aberdeen 
Proving Ground, 214 a large number of them were fired 
for the Army Ordnance Department at Purdue Uni¬ 
versity, 319 and some by the British Barrel Life 
Panel. 406 ’ 414-418 

Accuracy-Life 

General Statement. A caliber .50 aircraft machine 
gun barrel is ordinarily considered to have adequate 
accuracy if most of the bullets of a burst do not have 
appreciable yaw. Accuracy-life, therefore, is deter¬ 
mined by the point at which ‘ ‘serious keyholing” begins. 

The outstanding characteristic of the nitrided, 
chromium-plated barrels with choked muzzles was 
their amazingly long accuracy-life. The barrel could 
be fired until it was red hot on the outside and until 
the bore from the origin to a foot or so from the muz¬ 
zle was severely eroded, and still the barrel fired 
accurately—often firing more accurately toward the 
end of a long firing program than at the beginning. 



6 


Figure 11 . Photographs of representative caliber .50 
AP-M2 bullets collected in sawdust to show variation in 
engraving. (Figure 9 in NDRC Report A-409.) 


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466 


NITRIDED AND CHROMIUM-PLATED MACHINE GUN BARRELS 


Almost as interesting as the long accuracy-life was 
the consistency of performance of the barrels. 

Barrels Fired on Mild Schedule. Barrel No. 49 was 
the first successful barrel of this type. * 1 It was fired on 
the 5 X 50 (1) schedule approximately 4,000 rounds 
before it was decided to stop the test. The accuracy 
of the barrel was better on the last firing than on the 
first. In comparison, the accuracy of a standard bar¬ 
rel, such as No. 58, broke down at about 180 rounds 
during the firing of the first group, and at 90 rounds 
during the firing of the second group. Representative 
targets of barrels 49 and 58 are shown as follows: 
Figure 5 is in the third group 1 and Figure 6 in the 
eleventh and twelfth groups fired in barrel 49. Figure 
7 represents the first group, and Figure 8 the second 
group fired in barrel 58. 

Keyholing bullets are those which are unstable be¬ 
cause of improper engaging of the lands. This occurs 
if the land diameter is too great or if the longitudinal 
profile of the lands has the wrong shape. The land 
diameter increases permanently by erosion and swag¬ 
ing, and temporarily by thermal expansion as the 
barrel heats up. 

Erosion and swaging were minimized in these bar¬ 
rels by chromium-plating and nitriding, respectively. 
In Figure 12 are shown star gauge curves of a barrel 
nitrided and chromium-plated (No. 49), a barrel just 
chromium-plated (No. 61), and a standard barrel 
(No. 58), all fired the 5 X 50 (1) schedule. After 
one group, the land diameter of barrel No. 58 was in¬ 
creased throughout its length. After eight groups, the 
land diameter of barrel No. 49 was not increased ex¬ 
cept in the region where the chromium had been re¬ 
moved (breechward of 16 in. from the breech). 
Furthermore, after firing, the lands in barrel No. 49 
reached their full height abruptly, whereas those of 
No. 58 rose to approximately full height over a long 
section. An abrupt rise of the lands promotes good 
engraving of the jacket of the bullet. 

Barrels Fired on the Usual More Severe Testing Sched¬ 
ules. In Figures 9 and 10 are shown photographs of typi¬ 
cal targets obtained on firing the 5 X 100 (2) schedule. 
The bullets fired from the standard barrel, No. 237, 109 
began keyholing in the second 100-round burst, while 
the nitrided, chromium-plated barrel with choked 


1 Detailed data on barrels No. 49 and No. 58 may be found 
elsewhere. 109 

1 Warm-up and velocity rounds account for the fact that the 
total rounds shown on the target photographs are not divisible 
by 250. 


muzzle, No. JN3, fired no keyholing rounds in the 
ten bursts fired. 

At the Purdue University range of the Army Ord¬ 
nance Department, the nitrided, chromium-plated 
barrels with choked muzzles were fired with “combat 
ammunition” on the B-l and C-l schedules to 100 per 
cent keyholing 319 and compared with the standard 
barrel. The results are shown in Table 1. 


Table 1. Performance of caliber .50 aircraft barrels 
fired B-l and C-l schedules at Purdue University, using 
combat ammunition.* 


Type of barrel 

No. of 
barrels 
tested 

Type 

of 

test 

Average 
rounds to 
100 per cent 
keyholing 

Standard steel (D28272) f 

10 

C-l 

167 


5 

B-l 

230 

Chromium-plated, with choked 

14 

C-l 

293 

muzzle, not nitrided 

14 

B-l 

507 

Nitrided, chromium-plated, with 

5 

C-l 

319 

choked muzzle (D7162011)t 

5 

B-l 

1098 

Stellite-lined near breech 

6 

C-l 

295 

(D7161580)f 

5 

B-l 

455 


* 2API-M8, 2 I-Ml, and 1 Tr-MlO in each five rounds, 
t Army drawing number. 


Advantage of Choked Muzzle. The effect of thermal 
expansion in reducing the ability of the lands to en¬ 
grave the bullet adequately was minimized by the 
choked muzzle. If a barrel is choked to a diameter of 
0.494 in., a temperature of about 800 C is required to 
expand this bore to the nominal diameter of a caliber 
.50 barrel—0.500 in. Choking the bore of a caliber .50 
barrel over a length of only about a foot at the muz¬ 
zle is sufficient. A caliber .50 AP-M2 bullet can en¬ 
gage abruptly rising, full-height lands, without seri¬ 
ous slipping, when traveling almost at muzzle veloc¬ 
ity. Barrels were tried whose bore diameter over the 
full length was reduced to 0.495 in. by chromium 
plating. They had no better accuracy-life than those 
choked only near the muzzle. There was no particular 
advantage in the taper, as such, but it was a conven¬ 
ient way to get a choked muzzle and at the same 
time a large enough bore at the breech to prevent the 
development of excessive pressures. 

Velocity-Life 

General Statement. The Army Air Corps uses a drop 
of 200 fps in cold velocity as a criterion for rejecting a 


CONFIDENTIAL 








THE NITRIDED, CHROMIUM-PLATED AIRCRAFT BARREL 


467 


worn barrel. The number of rounds a barrel can fire 
on a given schedule before its cold velocity has 
dropped 200 fps is therefore considered its velocity- 
life. 


If the initial velocity is, to be maintained, there 
must be no decrease in the'resistance of the bullets 
to move which would result in inefficient burning of 
powder and in loWer starting pressures. In a chromi- 


- BEFORE FIRING 

- AFTER TWO 250-RD GROUPS 

- AFTER FOUR 250-RD GROUPS 


AFTER SIX 250-RD GROUPS 
AFTER EIGHT 250-RD GROUPS 





DISTANCE FROM BREECH IN INCHES 

Figure 12. Star gauge curves of caliber .50 aircraft barrels fired one or more groups on the 5 X 50 (1) schedule. (This 
figure had accompanied the manuscript of NDRC Report A-409 but was not reproduced with that report.) 


CONFIDENTIAL 




































468 


NITRIDED AND CHROMIUM-PLATED M ACHINE GUN B ARRELS 


um-plated, caliber .50 aircraft barrel, the starting 
pressure was essentially maintained as long as the 
chromium remained on the grooves at the origin of 
bore, for as the chromium was removed from the 
lands at this point, constriction of the grooves and 
swaging of the lands tended to produce a smooth bore 
with cold diameter of 0.505 in., and this constriction 
offered sufficient resistance to the start of the bullet 
so that the cold velocity dropped very little or actu- 

SCHEDULE 
• 5X100(2) 

■ 10X40(1) 

A 1X300 

□ 1X150 

O 1X100 

A 20X25-3 

x CONTINUOUS TO (00 PER CENT KEYHOLING 

USING COMBAT AMMUNITION 



ROUNDS FIRED 

Figure 13. Average velocity of nitrided, chromium- 
plated caliber .50 aircraft barrels with choked muzzles 
plotted against rounds fired according to seven different 
schedules. (This figure had accompanied the manu¬ 
script of NDRC Report A-409 but was not reproduced 
with that report.) 

ally increased. k Eventually the cold velocity dropped 
as the chromium was progressively removed from the 
grooves, the removal starting at the origin of bore 
and progressing muzzle ward. 

Velocity-Life of Nitrided , Plated Barrels. The veloc¬ 
ity of the nitrided, chromium-plated barrels in¬ 
creased at first as the grooves constricted and caused 
slightly higher gas pressures; then as the chromium 
was gradually removed from the vicinity of the origin 
of rifling and the steel eroded, the velocity fell off 

k A fuller discussion of the causes of velocity changes in 
barrels may be found elsewhere. 110 


rather regularly. In Figure 13 is shown a plot of veloc¬ 
ity versus rounds fired on seven different schedules. 
For a mild schedule such as 1 X 100 or 20 X 25 (3), 
the velocity-life was not much different from that of a 
standard barrel. On the more severe schedules, how¬ 
ever, the velocity-life was much better than that of a 
standard barrel. For example, on firing one 500-round 
group on the 5 X 100 (2) schedule, the velocity of a 
standard barrel dropped 400 to 500 fps as compared 
with a drop of about 100 fps for the nitrided, chro¬ 
mium-plated barrel. 

Gauge for Indicating Rejection Point of Barrels. No 
matter how good a barrel is, it will reach a point, if 
fired long enough, at which it should be discarded. An 
important problem with all barrels is the finding of a 
suitable field method for determining this rejection 
point. In general, caliber .50 aircraft barrels should 
be rejected if on the next short burst they would be 
firing bullets with a large yaw or the velocity drop 
would be more than 200 fps. The problem of a gauge 
for indicating the rejection point of a nitrided, chromi¬ 
um-plated barrel with a choked muzzle is resolved 
into a single problem of a method for indicating veloc¬ 
ity-drop, since the accuracy-life is much longer than 
the velocity-life. 

A gauge which is a very good indicator of the veloc¬ 
ity-drop of this barrel was developed. A drawing of it 
is shown in Figure 14, and a curve of velocity-drop 
versus gauge-advance is shown in Figure 15. 

The gauge is of the plug type, described in Section 
10.2.2, with a head 0.507 in. in diameter. It is entered 
from the breech end of a barrel, and the amount of its 
forward run beyond the zero point is observed. Then 
the velocity-drop of the barrel can be read from Fig¬ 
ure 15. 

Cause of Early Appearance of “Tippers” 
in some Barrels 

After the nitrided, chromium-plated barrels with 
choked muzzles were in production (800 barrels a 
day) accuracy acceptance tests showed that some 
were being manufactured that had a poor accuracy. 
In the acceptance test a 300-round burst of AP-M2 
ammunition was fired. Bullets with a yaw of as much 
as 20 degrees (tippers) were appearing in the first few 
rounds. The barrels, however, did not fire keyholing 
rounds at an early stage, nor did the point of 100 per 
cent keyholes appear earlier than usual. 

Nothing could be found wrong with the barrels on 
routine inspection. After a detailed investigation it 


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THE NITRIDED, CHROMIUM-PLATED AIRCRAFT BARREL 


469 


was found that the poor performance was caused by 
roughness of the bore. 110 The roughness was in the 
originally machined steel surface, which remained 
rough after electropolishing. The plating, if anything, 
accentuated the roughness. 

Another possible explanation of the poorer accu¬ 
racy of some plated barrels over others is in the differ¬ 
ence in the methods used in manufacturing the bar¬ 


rels. It was shown that chrpmium-plated barrels and 
stellite-lined barrels which had a separate muzzle 
bearing had greater dispersion, because of a higher 
percentage of tippers, than the same types of barrels 
with an integral muzzle bearing. Several manufactur¬ 
ing procedures were varied in both cases so it was not 
possible to determine what factor caused the differ¬ 
ence in behavior. 




ONE-TOOL STEEL/ TS ONE-PIN STEEL / T\ 

FINISH f HARDEN^ 3/32 X 1/2 DRIVEN^ 


TOLERANCE OF ± 1/64 ALLOWED ON ALL 
DIMENSIONS UNLESS OTHERWISE SPECIFIED 

Figure 14. Drawing of 0.507-in. plug gauge for determining rejection point of nitrided, chromium-plated, caliber .50 
aircraft barrels with choked muzzles. (Figure 21 in NDRC Report A-409.) 


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470 


NITRIDED AND CHROMIUM-PLATED MACHINE GUN BARRELS 



ADVANCE OF 0.507 IN. GAUGE IN INCHES 


Figure 15. Curve showing drop in muzzle velocity cor¬ 
responding to advance of the 0.507-in. plug gauge in 
eroded, nitrided, chromium-plated, caliber .50 aircraft 
barrels with choked muzzles. (Figure 22 in NDRC 
Report A-490.) 

23 2 OTHER MODIFICATIONS OF MACHINE 
GUN BARRELS 

23 2 1 Caliber .50 Aircraft Barrels 

Chromium-Plated Barrels Not Nitrided 

In Section 23.1 was discussed the nitrided, chro¬ 
mium-plated caliber .50 aircraft barrel with choked 
muzzle. A few words should be said about barrels 
identical to these except that they were not nitrided. 
In addition, the steel to be replaced by chromium was 
removed from some of these nonhardened barrels by 
machining rather than by electropolishing. These 
barrels were made before the facilities for large-scale 
nit riding had been found. 

Even without the nitriding, the chromium-plated 
barrel with choked muzzle was much superior in 
accuracy-life to the standard barrel, and in fact was 
as good in this respect as the barrel containing a 
stellite liner (Chapter 22), according to firing tests. 
The data are recorded in Table 1. Although superior 
to standard barrels, the nonhardened barrels, whether 
machined or electropolished prior to plating, were not 
as good with respect either to accuracy-life or to 
velocity-life as similar barrels that were nitrided; 
therefore, nitriding was adopted and facilities for 
large-scale nitriding were easily found. 

• 

Chromium-Plated Barrels of Other Steels 
with Hardened Bores 

The improved caliber .50 aircraft barrels described 
in Section 23.1 were made of WD 4150 steel. Re¬ 
search on special steels was carried out, as described 


in Section 16.4.11, to find an improved type of steel 
for machine gun barrels. For barrels that were to be 
plated with an erosion-resistant material such as 
chromium, it was not as important to have a steel 
that resisted attack by the powder gases as it was to 
have one with better high temperature properties 
than nitrided WD 4150 steel. The search for these 
superior steels was narrowed down to a few which 
would develop, upon hardening, a tough case that 
had good hardness at temperatures approaching 
800 C and good resistance to tempering. 

The following steels 1 were made into caliber .50 
aircraft barrels, which after bore-hardening were to 
be chromium-plated and tested: 

Nitralloy, obtained from A. Milne and Company, 
New York. Analysis: 0.43 C, 0.56 Mg, 0.017 P, 0.018 
S, 0.30 Si, 1.58 Cr, 0.35 Mo, 1.04 Al. 

Molybdenum Steel , obtained from Ford Motor 
Company. Analysis: 3 Mo, 0.13 C. 

Molybdenum Steel, made to order by Allegheny- 
Ludlum Steel Corporation. Analysis (average of 4 
heats): 0.33 C, 0.45 Mn, 0.20 P, 0.015 S, 0.29 Si, 
2.05 Mo, 0.10 Ni, 0.03 Cr. 

Chro-mow, obtained from the Crucible Steel Com¬ 
pany. Analysis: 5 Cr, 1.35 Mo, 1.25 W, 1 Si, 0.30 C. 

Potomac Die-Steel, obtained from the Allegheny- 
Ludlum Steel Corporation. Analysis: 5 Cr, 1.7 Mo, 
1.3 W, 1 Si, 0.33 C. 

Only the first three of these steels were tested as 
barrels. There was not time to conclude the work on 
barrels of the Chro-mow and Potomac hot-die steels, 
although barrels were made up and two of the Chro- 
mow barrels were nitrided. 

Nitrided Nitralloy has a very hard case, its hardness 
near the surface being of the order of 1000 VPN. The 
hot hardness of the case is excellent. Aircraft barrels 
of this steel were nitrided, electropolished, and chromi¬ 
um-plated, and subjected to firing tests. The lands 
near the origin of bore stood up very well. The ad¬ 
herence of the chromium to the steel and the resist¬ 
ance to swaging of the lands were better than in 
nitrided WD 4150 steel barrels. But the barrels failed 
near the muzzle end, where sections of lands broke 
out. The case is apparently too brittle for strains near 
the muzzle of this barrel. 

The bore-surface of barrels of both the molyb¬ 
denum steels were induction-hardened, electropol¬ 
ished, and chromium-plated. They performed better 
than chromium plated barrels of WD 4150 steel but 

1 Some of these same steels were made into barrels for stel¬ 
lite liners, as described in Section 24.5.3. 


CONFIDENTIAL 














OTHER MODIFICATIONS OF MACHINE GUN BARRELS 


471 


were not superior to those of WD 4150 steel nitrided 
and chromium-plated. 

Studies of ‘Tree-Run” Barrels 

As discussed in Section23.1.4, the nitrided, chromi¬ 
um-plated caliber .50 aircraft barrel with choked 
muzzle had a very long accuracy-life compared with 
its velocity-life. Such a barrel could be eroded to a 
smooth bore condition over the section extending 
from the origin of bore to within about a foot of the 
muzzle and still fire accurately, but with the barrel 
in this condition the muzzle velocity was several 
hundred feet per second lower than in a new barrel. 

The web of the powder was such that it burned 
efficiently and built up the proper pressure only when 
the bullet had an adequate starting pressure. When 
there were no lands at the origin of bore or no con¬ 
striction to resist movement of the bullet, the powder 
pressure did not reach the required high peak. 

In view of these considerations, it was thought that 
a barrel with a longer velocity-life and adequate 
accuracy-life would result, without giving excessive 
initial pressure, if a barrel were reamed to give a few 
inches of free run to the bullet. A powder which would 
burn fast enough to develop the required pressure in 
a smooth-bore barrel would have to be used. Possibly 
the barrel could then be eroded for a considerable 
distance ahead of the origin of bore without greatly 
affecting the muzzle velocity. 

Single-base powders having a smaller web than is 
usual for caliber .50 ammunition were loaded into 
cartridge cases and bullets crimped in. They were 
fired in steel barrels having various lengths of free run 
and chamber pressures were measured by means of a 
copper ball crusher gauge. A “free-run” of 1% in. 
and 244 grains of Ml powder of one particular ex¬ 
perimental lot were found to be about the optimum 
conditions. 109 ’ 110 

Using the components of AP-M2 ammunition, ex¬ 
cept for the powder, Frankford Arsenal then loaded 
10,000 rounds containing the lot of powder that was 
found best. Chromium-plated, free-run, aircraft bar¬ 
rels, with choked muzzles, were made. Some of the 
barrels were nitrided prior to plating. In the firing 
tests, using the 5 X 100 (2) schedule and the special 
ammunition, the velocity-drop was considerable, 
probably because of severe erosion in the free-run 
zone. 

It was thought that this erosion could be consider¬ 
ably lessened by experimenting to find the optimum 


diameter of the free-run cylinder and a better bore- 
surface coating. Because of (the fact that the project 
was never given a high priority, the contract termin¬ 
ated before this program could be rounded out. 

23.2.2 Nitrided, Chromium-Plated Caliber 

.50 Heavy Barrels With Choked 
Muzzles 

Although the first experiments with liners were 
made using the caliber .50 heavy barrel,” 1 research 
with the explicit object of improving this particular 
barrel was postponed until the major work on the 
aircraft barrel was completed. The standard mono¬ 
bloc steel, heavy barrel has a good life as is, but it is 
obvious that its life can be improved considerably by 
inserting a liner of Stellite No. 21 (Section 22.4) or 
properly chromium-plating the bore, or both. 

Nitrided, chromium-plated heavy barrels with 
choked muzzles were made. Both hard and soft chro¬ 
mium plates were applied. They w^ere fired according 
to the 5 X 100 (1) schedule. 110 It was found that 
modifying barrels in this manner would give a large 
increase in accuracy-life and, when this firing schedule 
was used, would increase somewhat the velocity-life. 
On the more mild schedules, the velocity-life would 
probably not be significantly improved. 

23.2.3 Caliber .30 and Caliber .60 Barrels 

General Statement 

Inasmuch as the problems connected with caliber 
.30 and caliber .60 machine gun barrels are very 
similar to those of the caliber .50 barrels, little further 
research was required to improve them. 

Nitrided and Chromium-Plated Caliber .30 
Aircraft Barrels with Choked Muzzles 

Caliber .30 aircraft barrels were tested in the air¬ 
craft gun which fires 1,200 to 1,300 rounds per min¬ 
ute. Two firing schedules were used with armor-pierc¬ 
ing ammunition: (1) A series of 300-round bursts 
with complete cooling after each burst—1 X 300 
schedule, and (2) three 300-round bursts with 2 min¬ 
utes between bursts—3 X 300 (2) schedule. 

The barrels which were tested were chromium- 
plated to give a choked muzzle of 0.292 + 0.004-in. 


See footnote (c) of Chapter 11. 


CONFIDENTIAL 





472 


NITRIDED AND CHROMIUM-PLATED MACHINE GUN BARRELS 


bore diameter. Some of these were nitrided prior to 
plating. 

The following conclusions were drawn from this 
investigation: (1) As in the case of the caliber .50 
barrels, the more severe the firing schedule, the 
greater was the advantage of nitriding and taper- 
chromium-plating the bore. When the 1 X 300 sched¬ 
ule was employed, the accuracy-life of the chromium- 
plated barrels with choked muzzles was considerably 
better than that of a standard barrel, and the ve¬ 
locity-life was about 25 per cent better. With the 
3 X 300 (2) schedule, both accuracy-life and velocity- 
life were much better for the modified than for the 
standard barrel. (2) The soft type of chromium ap¬ 
peared to be no better in this barrel than the hard 
chromium. Nitriding the bore, in addition to plating 
it, was advantageous. 

Nitrided and Chromium-Plated Caliber .60 
Barrels 

Caliber .60 barrels, a few of which had been ni¬ 
trided, were chromium-plated. It was particularly 
important in the chromium-plated barrels for this 


gun to have the cylindrical section lying between the 
forward end of the cartridge case and the run-out of 
rifling sufficiently large to eliminate the possibility of 
excessive pressures. Moreover, barrels performed in 
general better when the bore diameter near the origin 
was on the large side of the tolerance. Some barrels 
with choked muzzles were tested, but no conclusive 
results were obtained. 

The firing schedule used to test these barrels at 
Aberdeen Proving Ground 215 was supposed to con¬ 
sist of six 50-round bursts separated by 2 min inter¬ 
vals. Malfunctioning of the gun and ammunition, 
and the necessary replacement of parts during the 
test usually prevented this schedule from being fol¬ 
lowed. Consequently, it was difficult to compare 
barrels made in different ways. 

In general, it can be said that chromium-plating 
the bore of a caliber .60 barrel, with or without prior 
nitriding, could be expected to increase its life by a 
factor of at least 2 or 3 when fired on a schedule 
which would wear out the standard steel barrel in 
200 to 300 rounds. 110 Therefore, the investigation 
should be continued when a more reliable gun mech¬ 
anism is available. 


CONFIDENTIAL 



Chapter 24 

/ 

BARRELS BOTH STELLITE-LINED AND CHROMltjM-PLATED 

By J. F. Schairer a 


2 4 1 DEVELOPMENT AND PERFORMANCE 
OF "COMBINATION” CALIBER .50 
AIRCRAFT BARRELS 

W hile the two types of improved caliber .50 
aircraft machine gun barrels, stellite-lined (de¬ 
scribed in Chapter 22), and nitrided and chromium- 
plated (described in Chapter 23) were being devel¬ 
oped by Division 1, NDRC, at Crane Company and 
Geophysical Laboratory respectively, their advan¬ 
tages and disadvantages were noted. It soon became 
evident that a third and still better barrel should 
result from a combination of the stellite liner with 
choked-muzzle chromium plate on the steel bore 
ahead of the liner. Thereupon, the development of the 
stellite-lined and chromium-plated barrel with choked 
muzzle (or “combination” barrel) was initiated at Geo¬ 
physical Laboratory and developed by the coopera¬ 
tive efforts of this laboratory, the Crane Company, 
and the National Bureau of Standards. 

2411 Accuracy- and Velocity-Life of 
Improved Machine Gun Barrels 

Although the stellite-lined barrel showed a marked 
improvement in accuracy-life over steel barrels with 
no liner, as described in Section 22.2.2, the most 
outstanding characteristic of the lined barrel was the 
vast improvement in velocity-life. This long velocity- 
life is the result of the phenomenal resistance of the 
liner material to bore enlargement by either powder 
gas erosion (chemical attack or melting or both) or 
by the swaging impact of the bullets. In fact, a small 
constriction of both land and groove diameters oc¬ 
curs during the first burst of fire resulting in a ve¬ 
locity increase and during additional firing bore en¬ 
largement and velocity decrease only take place very 
slowly. 

When one of the more severe firing schedules is 
used (Section 23.1.3), the cold muzzle velocity will be 
little if any below the initial velocity and the liner is 
relatively new and unworn when the barrel is rejected 
for inaccuracy. The relatively short accuracjMife, as 

‘Special Assistant, Division 1, NDRC. (Present address: 
Geophysical Laboratory, Carnegie Institution of Washington.) 


compared to velocity-life, is largely the result of bore 
enlargement ahead of the liner by erosion of steel and 
of the thermal expansion of the barrel at the high 
temperatures (Section 5.6.3) attained in severe firing 
with the resultant failure of the bullets to engage the 
rifling properly and attain adequate spin for stable 
flight. 

In contrast, although nitrided and choked-muzzle, 
chromium-plated barrels showed (see Section 23.1.4) 
a marked improvement in velocity-life over non- 
plated steel barrels, their most outstanding charac¬ 
teristic was a vast improvement in accuracy-life. The 
long accuracy-life of this barrel is largely the result of 
the erosion-resistant choked-muzzle, which is re¬ 
sponsible for maintaining stability of bullets through¬ 
out long bursts (as already discussed more fully in 
Section 23.1.4), and barrels were rejected for exces¬ 
sive velocity drop long before they fired any inac¬ 
curate bullets. 

Thus, the performances of the two types of im¬ 
proved machine gun barrels, stellite-lined, and ni¬ 
trided and chromium-plated, were complementary, 
the former showing a very large improvement in 
velocity-life, the latter in accuracy-life, over un¬ 
modified steel barrels. 

24,1,2 Preparation and Testing of 
"Combination” Barrels 

The studies at the Geophysical Laboratory of the 
cause of failure of caliber .50 barrels made it clear 
that a barrel having a combination of a short breech 
liner of Stellite No. 21 and tapered-chromium-plate 
on the steel bore ahead of this liner should have a 
longer overall life than either the stellite-lined or 
nitrided and chromium-plated barrel alone. After 
this advantage had been demonstrated by the Geo¬ 
physical Laboratory a cooperative program for the 
preparation and testing of “combination” barrels was 
arranged between Geophysical Laboratory (and its 
subcontractor for plating, the Doehler-Jarvis Corpor¬ 
ation) and Crane Company. An additional advantage 
hoped for from the use of chromium plate ahead of 
the liner was a saving of strategic stellite by using a 
much shorter liner. 


CONFIDENTIAL 


473 



474 


BARRELS BOTH STELLITE-LINED AND CHROMIUM-PLATED 


Fourteen 10-lb caliber .50 aircraft barrels for each 
of three liner lengths were prepared with the steel 
bores chromium-plated ahead of investment-cast 
Stellite No. 21 liners of 4-in., 6-in., and 9-in. lengths. 
Machining and insertion was performed by Crane 
Company and the bore was electropolished and 
plated by the Doehler-Jarvis Corporation before liner 
insertion (Section 25.2.3). The procedure developed 
for the chromium plating of the bore ahead of a stel¬ 
lite liner was, with only slight modification, the same 
as used for full-length barrel plating. 

Four of each of the “combination” barrels with the 
liners of the various lengths were subjected to firing 
tests at Geophysical Laboratory 81 and two of each at 
Crane Company. 80 Concordant data were obtained 
and the remainder of the satisfactory barrels were 
delivered to the Ordnance Department for prelimi¬ 
nary tests. As a result of all these tests, 80 additional 
similar “combination” barrels with 9-in. length 
liners were delivered to the Ordnance Department 
for test. 

To evaluate the effects of bore hardening (by ni¬ 
triding) beneath the chromium plate ahead of 4-in., 
6-in., and 9-in. length stellite liners, two of each of 
these barrels were prepared for test by Division 1. 
Firing tests showed that bore hardening ahead of the 
liner had but little effect on performance and that 
this manufacturing complication was unnecessary. 81 

2413 Development of the 

CGL Firing Schedule 

During the development of the “combination” 
barrels just described and in the development of 
barrels showing even further improvements in per¬ 
formance (achieved by methods described later in 
this chapter in Sections 24.2, 24.3, and 24.5) it was 
necessary to increase the severity of the firing-test 
schedule. These caliber .50 aircraft machine gun bar¬ 
rels were so far superior in life and performance to 
anything previously known that a stepped-up firing 
schedule was required to give an adequate compar¬ 
ison of barrel performance and to indicate the full 
potentialities of these barrels. One requirement for 
such a schedule was that it yield a maximum of de¬ 
sired information on a barrel and that barrels could 
be fired to end of life in a reasonable time. Two con¬ 
tractors of Division 1 (Crane Company and Geo¬ 
physical Laboratory) collaborated in the develop¬ 
ment of the new firing-test schedule, hence the name 
“CGL” schedule. 


Several things were considered in the choosing of 
this schedule: 

1. The Air Forces required a barrel that is capable 
of firing a long continuous burst or bursts and closely 
spaced long intermittent bursts. 

2. Any schedule requiring the firing of a very long 
continuous burst or closely spaced long intermittent 
bursts which overheat and distort the barrel, because 
of the lack of high-temperature strength of gun steel, 
does not permit the evaluation of the relative merits 
of different liner materials. 

3. If too moderate a schedule is selected, it will re¬ 
quire a long firing test, besides not simulating ex¬ 
treme combat conditions. 

4. If possible, the schedule should be flexible, to 
provide for an evaluation of somewhat inferior ma¬ 
terials that might be used as substitutes, as well as 
for very superior materials and combinations. 

5. The same ammunition should be used in firing 
all special barrels to make the results comparable. 

Two schedules were seriously considered—one em¬ 
ploying a series of continuous bursts of arbitrarily 
fixed and diminishing length—the other a continuous 
burst of fixed length followed by a series of inter¬ 
mittent bursts. A schedule of the latter type was 
adopted as the firing-test schedule in the development 
of superior machine gun barrels and in the test of 
promising materials and combinations. This initial 
burst gives an indication of performance under the 
severe conditions of prolonged continuous fire and 
the subsequent repeated groups indicate perform¬ 
ance under the severe conditions of moderate length 
bursts at frequent intervals. 

The CGL schedule 80,81 calls for a long continuous 
burst of predetermined duration (usually 350 rounds) 
followed by 500-round groups fired in five bursts of 
100 rounds each with a 2-minute cooling interval 
between bursts, such groups being repeated until end 
of fife from inaccuracy (50-per cent keyholing bullets), 
excessive velocity drop (greater than 200 fps), exces¬ 
sive muzzle blast, or any dangerous condition of the 
barrel. After the long continuous burst and after each 
500-round group the barrel is cooled to room temper¬ 
ature. Accuracy during bursts is observed by means 
of a moving target. Cold- and hot-velocity measure¬ 
ments are taken at stated intervals. Figure 1 shows a 
data sheet for a caliber .50 aircraft barrel fired on the 
CGL-350 schedule. To provide a flexible base to eval¬ 
uate inferior barrels as well as very superior barrels, 
the length of the long continuous burst can be varied. 
A suffix thus (CGL-400, CGL-350, CGL-300, etc.) 


CONFIDENTIAL 




"combination” CALIBER .50 AIRCRAFT BARRELS 


475 


Data sheet for caliber .50 aircraft barrel No. AC155 fired on CGL-350 schedule. All targets at 1,000 in. Total 
rounds fired: erosion 1,350j all others 150. Instrumental velocity at 78 ft. (This figure is based/on one in progress report 
on Contract OEMsr-629.) \ 


Instr. vel. 20-rd cold accuracy 100-rd burst accuracy ^ Cyc]ic 

(fps) 90% of rd No. rds 95%ofrd No. rds bore rate, 

Test No. Cold Hot EHD* EVDf >20° yaw EHD* EVDf >20° yaw dia. rpm Remarks 


0. Initial 2640 5.5 

4.5 

0 

Continuous Burst Accuracy 

1. Cont. burst 

(350 rounds) 2650 2420 7.4 

3.6 

0 

First rd with yaw 20:337 
No. rounds with yaw 50: 0 
Number of yawers: 2 


Appearance satisfactory. 

Muzzle flash severe last 50 
rds. Mild erosion ahead of 
818 the liner; slight wear in 
liner; rifling clearly defined 
throughout. 


2. 500-rd group 


Burst 1 
Burst 2 
Burst 3 
Burst 4 


2625 

2600 

2465 

2360 

Burst 5 

2670 

2330 

500-rd group 
Burst 1 


2570 

Burst 2 


2360 

Burst 3 


2265 

Burst 4 


None 

Burst 5 

2415 

2030 


4.5 

4.5 

6.8 

10.8 

5.5 5.5 0 11.2 

9.3 

14 

10.2 

9 

10 10 20 11.3 


6.5 

0 

820 

6.7 

0 

820 

8.5 

10 

820 

10.5 

25 

Approx. 

820 

12.6 

100 

820 


9.7 

6 

Approx. 

780 

14.5 

36 

Approx. 

780 

14.3 

100 

Approx. 

780 

11 

100 

Approx. 

780 

11.8 

100 

780 


Muzzle flash severe during 
last 2 bursts. 

The barrel is eroded ahead of 
the liner for about 3 in. 
Rifling is clearly defined 
but worn in the center of 
the barrel. The liner is 
slightly worn. 

3 sec stoppage in burst 1 and 
2.5 sec stoppage in burst 
3. 


4. 500-rd group 
Burst 1 
Burst 2 
Burst 3 
Burst 4 
Burst 5 


* EHD = Extreme horizontal dispersion (in.). t EVD Extreme vertical dispersion (in.). 

Figure 1 


indicates the length of the long continuous burst 
which in all cases is followed by the regular 500-round 
groups of intermittent fire. All of the very superior 
barrels whose development is described in this chap¬ 
ter were tested on the CGL-350 schedule, using 
AP-M2 ammunition. The results are summarized 
later (Figure 6). 

2414 Performance of "Combination” Barrels 

Initial firing tests by Division 1 on “combination” 
caliber .50 aircraft barrels with 4-in., 6-in., and 9-in. 
length investment-cast liners of Stellite No. 21 and 
choked-muzzle chromium plate on the steel bore 
ahead of the liner were made on the continuous-to- 
keyholing (C—1) schedule b and on an intermittent 


b Defined in Chapter 22.3.3. 


(B—2) schedule. 0 The plated barrels with 4- and 6-in. 
liners were approximately equivalent in performance 
to a barrel with a 9-in. liner but no plate ahead of the 
liner. The plated barrel with 9-in. liner was far 
superior to either the regular stellite-lined barrel 
(Chapter 22) or the nitrided and chromium-plated 
barrel (Chapter 23) particularly on the intermittent 
schedule. The Air Corps wanted only the best barrels, 
so attention was concentrated in further tests on the 
“combination” barrel with a 9-in. liner with plate 
ahead of the liner. Besides the very severe CGL-350 
firing schedule used in further tests by Division 1, a 


c On the B—2 schedule a barrel is fired in 500-round groups 
consisting of five bursts of 100 rounds each with a 2-min 
cooling interval between bursts and complete cooling after 
each 500-round group. Such groups are repeated to end of life 
of the barrel. This schedule is designated “5 X 100 (2)” in 
Chapter 23. 


CONFIDENTIAL 


















476 


BARRELS BOTH STELLITE-LINED AND CHROMIUM-PLATED 


variety of schedules was used in tests by the Ordnance 
Department on barrels delivered by Division 1. 

As a result of the firing tests on very severe sched¬ 
ules one serious defect of “combination” barrels was 
discovered. Their long accuracy and velocity lives 
permitted firing of such duration and severity that 
many of the barrels failed in a dangerous manner by 
rupturing longitudinally near the muzzle end of the 
barrel. Such failure can be eliminated by reinforce¬ 
ment of the barrel, which is considered in the next 
section. 

24 2 EXTERNAL REINFORCING SLEEVE 
FOR BARRELS 

On a severe firing schedule, such as the CGL-350 
schedule, end of life of a “combination” barrel is 
brought about by a permanent expansion of the 
barrel in the region of the forward end of the liner. 
The sequence of events in the weakening of the barrel 
wall caused by the high temperatures during the fir¬ 
ing of long bursts is given in Section 5.6.4. Figure 16 
of Chapter 5 shows two caliber .50 aircraft barrels 
distorted during firing. It should be noted that dis¬ 
tortion was principally in one plane. Stellite-lined bar- 
els (with no plate ahead of the liner) will also blow 
out in the same manner if the firing of long bursts is 
continued long enough. Because of the much shorter 
accuracy life of this barrel as compared to the “com¬ 
bination” barrel, however, firing is usually stopped 
before the barrel can rupture. 

The expansion of various parts of the barrel was 
estimated 81 on the basis of star gauge data and out¬ 
side diameter measurements made before and after 
firing on a number of “combination” and other bar¬ 
rels. After firing the 350-round burst on the CGL-350 
schedule, “combination” barrels showed permanent 
expansion (cold) of the bore at 8 to 10 in. from the 
breech, that is, 2 to 4 in. behind the end of the liner. 
The average for seventeen barrels measured was 
0.0043 in. (range from 0.000 to 0.015 in.). Just ahead 
of the liner the expansion was about the same. For¬ 
ward of this the amount of expansion tapered off. 
Clearly, this expansion had to be prevented if the 
best use of either or both the liner and chromium 
plate was to be made. 

A simple method of reinforcing the barrel in this 
weak region was found to be highly successful, es¬ 
sentially eliminating the permanent expansion and 
greatly increasing the life of the “combination” bar¬ 
rel as well as making it a safe barrel. A thin cylindri¬ 


cal steel sleeve or collar, 6-in. long, was shrunk on the 
outside of the barrel, after it had been turned to a 
cylinder over the weakest section. It was located so 
that the region embraced by the forward 5 in. of the 
liner and 1 in. ahead of the liner was externally re¬ 
inforced. The weight of the sleeve had to be kept at 
a minimum because increase in barrel weight would 
have decreased the cyclic rate of fire of the gun. This 
matter is discussed more fully later in Section 24.3.2. 

The experiments on barrel reinforcement at Geo¬ 
physical Laboratory indicated that, (1) a reinforcing 
sleeve should start at about 7in. from the breech 
and extend to 133^ in. from the breech, (2) sleeves 
made of WD 4150 steel or of NE 8630 steel tubing are 
satisfactory, (3) wire winding of the barrel ahead of 
the sleeve is advantageous, but may not give enough 
improvement to justify the extra time, cost, and 
weight, and (4) external sleeves should be fixed in 
place, as by a spot weld at the forward end of the col¬ 
lar, to eliminate any possibility of their slipping. 
The experiments with “special contour” barrels de¬ 
scribed in Section 24.3.1 suggested that most of the 
benefit of an external reinforcing sleeve might be 
provided by machining the barrel to have outside di¬ 
mensions the same as those of a barrel with a shrunk- 
on external sleeve. Two alternative designs for “com¬ 
bination” barrels with integral reinforcement were 
prepared at Geophysical Laboratory and some bar¬ 
rels were made according to these designs which are 
designated as “GLA” and “GLB” in Figure 2. It was 
not possible to make the firing tests before termina¬ 
tion of the experimental program of Division 1. The 
barrels were turned over to the Ordnance Department 
with the recommendation that these tests be made. 

Firing tests 80 by Division 1, according to both the 
CGL-350 and CGL-400 schedules made on stellite- 
lined barrels without plate ahead of the liner but with 
reinforcing sleeves showed that such barrels gave 
considerably improved performance over regular 
stellite-lined barrels. Accordingly a few thousand 
such barrels for Service trials and use were to have 
been produced at Crane Company in the late summer 
of 1945 under an Ordnance Department contract. 
Substantial production of “combination” barrels 
with reinforcing sleeves also was planned. Although 
the expedient of adding reinforcing sleeves was satis¬ 
factory in principle for the utilization of barrels in 
stock, the machine gun barrel manufacturers pre¬ 
ferred to manufacture new barrels with integral steel 
reinforcement rather than to shrink-on external 
sleeves. 


CONFIDENTIAL 



CHANGE IN WEIGHT AND CONTOUR OF BARRELS 


477 




Figure 2. Designs for “combination” caliber .50 aircraft machine gun barrels with integral reinforcement at the 
forward end of the stellite liner. (Figure 24 in NDRC Report A-409.) 


The remarkable results of the studies Division 
1 on barrel reinforcement in general and other 
changes in external contour of barrels are described 
in the next section (Section 24.3). Regular and special 
caliber .50 aircraft barrels are pictured in Figure 3. 
The relative performance of all superior barrels in¬ 
cluding both “combination” barrels and stellite-lined 
barrels with an external reinforcing sleeve is shown 
later in Figure 6. Before new barrel forgings could be 
obtained the Army production contracts were can¬ 
celed immediately after V-J Day. 


24 3 CHANGE IN WEIGHT AND CONTOUR 
OF LINED AND "COMBINATION” 
BARRELS 

24,31 Effects of Barrel Weight 

on Performance 

Even before the external reinforcing sleeve had 
been added to stellite-lined barrels at the Geophysical 
Laboratory, a different approach to the same problem 
had been started at the Crane Company, another 
Divison 1 contractor. 



Figure 3. Photographs of regular and special caliber .50 aircraft barrels. (Figure 6 in NDRC Report A-408.) 


CONFIDENTIAL 





















































478 


BARRELS BOTH STELLITE-LINED AND CHROMIUM-PLATED 


To evaluate the effects of barrel weight on the per¬ 
formance of stellite-lined caliber .50 aircraft machine 
gun barrels, regular 9-in. liners of investment-cast 
Stellite No. 21 were inserted in two 13-lb caliber .50 
aircraft barrels (Figure 4B). This 13-lb barrel (36-in. 
long) was an obsolete barrel that had been replaced 
for Service use by the regular 10-lb barrel having the 
same length (Figure 4A.) 

Firing tests on the stellite-lined 13-lb barrels on the 
severe CGL-350 schedule showed a remarkable in¬ 
crease in life. Whereas the life of the regular 10-lb 
stellite-lined barrel on this schedule averages about 
750 rounds, the lined 13-lb barrels showed a life of 
about 3,500 rounds. In both cases rejection was 
caused by inaccuracy while velocity drop had been 
very slight. In contrast, a change in weight from 10 lb 
to 13 lb results in only a very slight increase in life 
with regular steel barrels without the liner, the total 
life on the same schedule being 170 and 220 rounds, 
respectively. 

24.3.2 Effects of Barrel Weight on 

Cyclic Rate of Gun 

Because of the Air Forces’ requirement for a high 
cyclic rate of fire in aircraft combat, it is important to 
note the effects of barrel weight on the cyclic rate of 
fire of aircraft machine guns. When a 13-lb barrel was 
fired at Crane Company in the Browning aircraft 
machine gun, M2, the cyclic rate was only 700 to 800 
rpm even with all of the oil removed from the buffer, 
whereas the cyclic rate with a 10-lb barrel in this gun 
can be adjusted to approximately 750 rpm with oil in 
the buffer and to about 1,000 rpm with the oil removed. 

Firing tests were made by the Ordnance Depart¬ 
ment on a series of barrels of different weights sup¬ 
plied by Division 1, using the newly adopted high¬ 
speed gun (M3). This gun has a rated cyclic rate of 
1,250 rpm with a 10-lb barrel. In the test firings in 
which all barrels were fired in the same gun, the 13-lb 
barrel reduced the rate by 200 rpm and an 11.75-lb 
barrel reduced it by only 70 rpm. These data sug¬ 
gested that, unless delays could be tolerated while the 
gun mechanism was modified, efforts should be con¬ 
centrated on the development of the best barrel with 
a weight not much in excess of 11.5 lb. 

24.3.3 Effects of Distribution of 
Weight on Performance 

The remarkable results of firing tests on stellite- 


lined 13-lb caliber .50 aircraft barrels were described 
in Section 24.3.1. Similar results were obtained on the 
CGL-350 schedule on special contour 13-lb barrels of 
design suggested by the Ordnance Department. These 
special contour barrels had a heavy cylindrical sec¬ 
tion extending from the breech end of the barrel to a 
point somewhat beyond the liner. The barrels were 
prepared at Springfield Armory and subjected to fir¬ 
ing tests at Crane Company as a part of a broad 
study of the effects of distribution of weight (changes 
in exterior contour of barrels) on the performance of 
stellite-lined and “combination” barrels. 

In view of the limitations imposed on barrel weight 
by the effects of increased weight on cyclic rate of fire of 
the gun (see Section 24.3.2), Crane Company pre¬ 
pared and tested barrels the weight of which was 
reduced by machining steel from 13-lb barrels to de¬ 
crease the diameter in selected areas. All of these 
tests were conducted with the M2 gun (no M3 guns 
were available at this time) at a cyclic rate of 750 
rpm. Two lined barrels were prepared for each of 8 
special designs shown in Figure 4. 

Another design (not shown in Figure 4) was similar in 
external contour to the barrel with a 6-in. external 
reinforcing sleeve (Figures 3B and 4J) except that 
the section of increased diameter was integral with 
the barrel and the weight was 10.9 lb. The perform¬ 
ance of the barrel with an integral sleeve was some¬ 
what superior to that of the barrel with the shrunk- 
on sleeve. The best performance per unit of weight 
was obtained from lined 13-lb barrels into which cir¬ 
cumferential grooves had been cut so that the con¬ 
tour at the root of the groove was the same as that of 
the standard 10-lb barrel. The result was a barrel 
(Figures 3D and 4C) having integral, transverse fins, 
with a weight of about 11.7 lb. The performance 
of this barrel was equivalent to that of the lined 
13-lb barrel without fins (Figure 4B) and represents 
an outstanding improvement in performance over 
stellite-lined 10-lb barrels with only a very slight re¬ 
duction in the cyclic rate of fire. Such fins provide a 
maximum of barrel reinforcement and cooling surface 
with a minimum of additional barrel weight. 

The two designs described in the preceding para¬ 
graph were selected for test with choked-muzzle chro¬ 
mium plating ahead of the liner. Two barrels of each 
design were prepared. The performance of the plated 
barrel with the integral reinforcing sleeve (weight 10.9 
lb) was equivalent to that of the lined 13-lb barrel de¬ 
scribed above with fins (weight 11.7 lb) but with no 
plate ahead of the liner. 


CONFIDENTIAL 



CHANGE IN WEIGHT AND CONTOUR OF BARRELS 


479 


'«5 = ro 

is. O- 

a? 


36.000_o3o" 


9 “ APPROX- - -* ” “ ““.U 5 U I q 3 q“ 

4- r i£- 3 /ig"r -3% [pO NOT PRODUCE SHARP GROOVE ^INCLUDED TAPER PER INCH =.011 APPROX -.015“ 


O o 

r- q- 

i£) r 


V 7 ioo“ [P 2Q "-^H^ 


^INCLUDED TAPER PER INCH =.051" (0-35348A8) 

-32.000" TO INTERSECTION OF TAPER-- 


I 




REGULAR 10.2 LB AC BARREL 
-k WT 10.2 LB 


.789"4^*p 

+J0 °" 1*2.920"-.040* 




1500" 


58 9 


1.567]'DIA 1567" DIA 1567" DIA 1.545" DIA ^??« 0 j. 0 A 30 " 1365" DIA L278" DIA 1.182" DIA | L045 '? 1 ^" ^29“ 

-*-1-A-A_ IA56I DIA 1 _ t 11.098 DIA-^H F-^32-■ 




“2 o 
Q 


(D-35348-3) 


1.567 DIA 1.498 Dl, 


f f f |.|40" DIA^ AFTER CHROMIUM PLATING—^ 

IA 1.410" DIA 1.315" DIA 1 1.230" DIA * REGULAR 13 LB AC BARREL 

WT 13.16 LB 


> 5“ ai5“ 1567" DIA 1.567" DIA L567" DIA 1.545" DIA 3&000 030 » 





1 I ! I ! I.lio" DIA k 1-021" DIA t ' 

1.480 DIA L445 DIA 1.360 DIA 1.280 DIA 1.200 DIA SPECIAL CONTOUR (UNFINNED) 
-32.000"-.030" TO INTERSECTION OF TAPERS----- *\ WT 12.96 LB 



i", ‘• 03 ° U ,, 1.315" DIA 1237" DIA U50" DIA i* .f-t-2" 

1.473 DIA L360 DIA I 1.280 DIA |.200“ DIA I 


.789"+-4 
+.ioo"|—2.920.; 


ROOT DIA OF GROOVES TO BE SAME AS OUTu . . ' „ I. ak * a 3 * 

SIDE DIA OF IQ2 LB CAL .50 AC MG BARREL L445 DIA ^1.400 DIA A, 

-32.000" TO INTERSECTION OF TAPER 



\ if— kk? 

U072 DIA H. 


o J. 
- 2 2 - 

----ZJfz- 

J-- 

Uio" DIA-J M.072" DIA H.OOO" Dia 

SPECIAL CONTOUR (FINNED) 
WT 11.66 LB 






36.000“^ - 

I.O: 

30" 

f 

-T r 

1 £ 


- - 


1.570" 


-. 010 “ 


NOTE- DIMENSIONS SHOWING DIAMETERS IN DIAGRAMS 
"B", "C", "D" AND ”E" TAKEN AT 11 / 2 “ INTERVALS 


SPECIAL CAL .50 AIRCRAFT BARREL WITH 
STRAIGHT TAPER. WT 11.6 LB 


-14.000"- 




-36.000' 




-JO 30" 

•INCL TAPER PER INCH = .Oil" APPROX 


1.030 


1.450 


SPECIAL CAL .50 AIRCRAFT 8 ARREL WITH 1.450" OD 
BREECH SECTION. WT 10.97 LB 


H S 


8 - 

I 


■t 

PS 


*ZoiO"l“ = "'2.92 15 —4gl*- 


+-T- 


-94- 


o'g 

3?" 




45 TAPER 




-||*"- 


jS 


-4- 

dfc: 


36.000 _ 030 .. 


0° 34' TAPER 




- 2 i- 






= oV? 

8 ? 


i |7 


r - „ »1 *-0 




M/I 6 "R -1/16" 

SPECIAL CAL .50 AIRCRAFT BARREL WITH DOUBLE TAPER 
AND CENTER BEARING. WT 11.67 LB 



ROOT DIA OF GROOVES TO BE SAME AS OUT- 

- 0° 34' TAPER^l SIDE DIAM OF, l(X2 LB CAL 50 AC MG BARREL 89 


4*- 


Ml ® ^ - 1 / 16 “ i 


,• VI© n -1/16 

Vlfc SPECIAL CAL .50 AIRCRAFT BARREL WITH DOUBLE TAPER AND 
CENTER BEARING, FINNED AHEAD. WT IL28 LB 


TURN L2II" INCL TAPER PER WCH = . 0 I|" APPROX 


1.030" 


i 


INCL TAPER PER INCH = .051" 


L209" DIA OF BORE IN COLLAR 
FOR SHRINK FIT 


CAL .50 AIRCRAFT BARREL WITH 6 " REINFORCING SLEEVE 
WT IQ.75 LB 


Figure 4. Eight designs of caliber .50 aircraft machine gun barrels having special contours. (Drawing submitted by 
Crane Company with interim report on Contract OEMsr-629.) 


CONFIDENTIAL 





































































































































































































































































































































































































































































































































































































































































































480 


BARRELS BOTH STELLITE-LINED AND CHROMIUM-PLATED 


The finned barrel with both liner and plate (weight 
11.7 lb) was far superior to any other aircraft barrel 
tested and showed a life of about 5,000 rounds on the 
severe CGL-350 schedule. This barrel life approx¬ 
imately is equal to the life of the gun mechanism and 
both gun and barrel can be discarded as a unit. One of 
the lined and plated finned barrels was rejected after 
5,000 rounds because of a 200-fps velocity drop. The 
other burst near the muzzle after more than 5,000 
rounds had been fired. The use of a barrel steel having 
better high temperature properties than the WD 4150 
steel regularly used for machine gun barrels would 
prevent such rupture. (This question of choice of bar¬ 
rel steel is discussed later in Section 24.5.) 

In the stellite-lined and chromium-plated finned 
barrel, the full usefulness of the erosion resistant liner 
is attained and liner and barrel wear out at about the 
same time. Preparations for procurement of the neces¬ 
sary barrel forgings for the production of these very 
superior barrels were in progress by the Ordnance 
Department when the procurement program was cur¬ 
tailed after V-J Day. 

24 3 4 Aluminum-Clad Barrels 

In order to evaluate the effects on performance that 
might be achieved by the external cladding of regular 
steel barrels with a lightweight metal of high thermal 
conductivity, regular 10-lb aircraft barrels were clad 
at the Al-Fin Corporation with aluminum muffs by 
the Al-Fin Thermit process. 559 These barrels were ma¬ 
chined at Crane Company to yield a barrel (Figure 5) 
with circumferential fins, and stellite liners were in¬ 
serted in some of these barrels. The total barrel weight 
was 11.5 lb. Firing tests on unlined barrels showed 
some improvement in performance over 10-lb steel 
barrels with no aluminum muff but severe powder gas 
erosion and flattening of the lands by the swaging im¬ 
pact of the projectile terminated life at an early stage. 
When the stellite-lined, aluminum clad, finned bar¬ 
rels were tested on the CGL-350 schedule, the per¬ 
formance was better than that of stellite-lined barrels 
without muffs, but the aluminum melted and dripped 
from the hot barrel. 

Thereupon, two more lined barrels with solid alum¬ 
inum muffs were encased in a thin steel tube to hold 
in place any aluminum which might melt and to take 
advantage of the heat of fusion of aluminum to cool 
the barrel. These barrels, which weighed about 13.51b, 
showed a performance somewhat superior to that of 
the unplated and lined 13-lb steel barrels described in 


Section 24.3.1, but the weight was so great as to make 
this barrel impractical. 

Although it might have been possible to obtain 
comparable performance with reduced weight more 
suitably distributed, it was felt that lined and chro- 
^mium-plated barrels with integral steel fins were more 
promising for immediate Service application. There¬ 
fore, no further work was done on aluminum clad 
barrels. Some preliminary tests on copper clad bar¬ 
rels were made by the Ordnance Department as a re¬ 
sult of these studies by Division 1, NDRC, on alum¬ 
inum clad barrels. 


24 4 PERFORMANCE OF VARIOUS 
IMPROVED CALIBER .50 AIRCRAFT 
BARRELS 

In order to show clearly the remarkable degree of 
improvement in firing performance of caliber .50 air¬ 
craft machine gun barrels that it has been possible to 
make, the pertinent data are summarized in Figure 6, 
where comparison is made on the very severe CGL- 
350 schedule already described in Section 24 1.3. All 
firings were made with AP-M2 ammunition and all 
barrels were made of regular WD 4150 machine gun 
steel. In each case the number of rounds of total life 
is a representative value. In most cases it is an aver¬ 
age from firings on ten or more barrels. In all cases at 
least two barrels were fired and gave closely agreeing 
test data. 

24 5 CHANGE OF BARREL STEEL TO 
ENHANCE PERFORMANCE 

24 5 1 Barrel Heating During Firing 

The high overall barrel temperatures attained dur¬ 
ing severe firing (Section 5.4.2) limit the life and per¬ 
formance of machine gun barrels in terms of both the 
length of continuous bursts and the length and fre¬ 
quency of intermittent bursts. This situation is not 
corrected by the use of an erosion resistant liner or 
plate, and therefore it prevents their complete utili¬ 
zation. 

Temperatures as high as 900 C at a point 4.35 in. 
from the breech and 0.068 in. from the bore of a 10-lb 
caliber .50 aircraft barrel were recorded during the 
firing of severe schedules, as described in Section 
5.5.1. A temporary increase in bore diameter results 
from the thermal expansion of the barrel steel at these 


CONFIDENTIAL 



CHANGE OF BARREL STEEL TO ENHANCE PERFORMANCE 


481 


high temperatures and such expansion tends to pre¬ 
vent engagement of the bullets in the rifling and the 
attainment of proper spin for stable flight. Fortu¬ 
nately, erosion-resistant choked-muzzle chromium 
plating overcomes this difficulty in large part. 

The lack of high-temperature strength and hard¬ 
ness of the barrel steel is a more serious limitation. A 
barrel that distorts or “wilts” cannot be expected to 
fire accurately and safely. This difficult}^ can be miti¬ 
gated in lined and plated barrels by (1) changing the 
weight (amount) and distribution (contour 1 ) of steel 
in a barrel, (2) by cooling (internal or external), and 
(3) by a change in barrel steel composition. The use 
of the first of these methods to secure vastly enhanced 
performance was just described in Section 24.3. The 
second method is discussed in Section 5.7. 

There is good evidence from firing tests summarized 
later in this chapter that moderate to substantial im¬ 
provement in performance can be achieved in machine 
gun barrels with a stellite liner, with or without 
choked-muzzle chromium plate ahead of the liner, by 



Figure 5. Aluminum clad caliber .50 machine gun 
barrel with circumferential fins and a regular barrel 
for comparison. (This figure was taken from a progress 
report by Crane Company on Contract OEMsr-629.) 

the use of a barrel steel having better high-tempera¬ 
ture strength and short-time tensile properties than 
ordinary WD 4150 machine gun barrel steel to pre¬ 
vent permanent expansion. 


10 II 12 13 lb 

iiiiiiiiiimiiiiii 


NUMBER OF ROUNDS'. 0 IOOO 2000 3000 4000 5000 


mnm 


Wt 10.2 lb 


a 




Wt 13 lb 


170 


220 


Wt 10.2 lb 



Wt 10.2 lb 




wt 10.9 lb 




Wt 10.9 lb 




Wt 13 lb 



Wt M.7 lb 


750 


750 


1350 


1350 


3500 


3500 


3500 


5000 


LEGEND ill STELLITE LINER WM NITRIDED 8 CHROMIUM PLATED CHROMIUM PLATED 


REINFORCING SLEEVE ""IIZ INTEGRAL CIRCUMFERENTIAL FINS 


Figure 6. The increase in life of caliber .50 machine gun barrels that has resulted from the use of a stellite liner, 
chromium plating, and the combination of these two modifications. 



CONFIDENTIAL 




























































































482 


BARRELS BOTH STELLITE-LINED AND CHROMIUM-PLATED 


24 5 2 Special Alloy Steels for Gun Barrels 

The possible use of various steels and special high- 
iron alloys as bore-surface materials has already been 
discussed in Section 16.4.11. The conclusion was 
reached that because of their lack of resistance to 
powder gas erosion, no steels or high-iron alloys tested 
have shown any outstanding promise as bore-surface 
materials under severe firing conditions. Trials of 
special barrel steels as a base for choked-muzzle chro¬ 
mium plate on a hardened steel bore and of methods 
of hardening steel bores before plating have already 
been described in Section 23.2.1. These steels were 
selected for their good hot-hardness and low rate of 
loss of hardness with time when heated in the range 
of temperatures attained by caliber .50 aircraft bar¬ 
rels during severe firing schedules. Of the steels tested, 
only barrels of molybdenum steel (Mo, 3.0%; C, 
0.13%), whose bores had been induction hardened 
and then plated with choked-muzzle chromium plate, 
showed performance equal to or slightly better than ni- 
trided and similarly plated barrels of regular WD 4150 
machine gun barrel steel. 

In contrast, the use of a barrel steel with better 
high-temperature properties than WD 4150 gun steel 
gave moderate to substantial improvements in per¬ 
formance when machine gun barrels with a stellite 
liner, with or without choked-muzzle chromium plate 
ahead of the liner, were subjected to firing tests. Be¬ 
fore termination of the experimental program of Divi¬ 
sion 1 there was time for only a reconnaissance of the 
effects of change of composition and properties of 
barrel steel on the performance of stellite-lined and 
chromium-plated machine gun barrels. Enough spe¬ 
cial steel barrels were made and firing tests completed 
to indicate that by change of barrel steel at least a 
doubling of life could be obtained on severe schedules 
and that longer bursts of fire were possible and feas¬ 
ible without serious impairment of barrel life. Under 
the stress of war conditions, only a few readily avail¬ 
able special steels that could be obtained as bar stock 
or forgings of the proper size and which were ma¬ 
chinable by conventional methods could be tested 
as barrels. d It is strongly recommended that more 
systematic and extensive studies of the effect of bar- 


d The Ordnance Department was very much interested in 
these reconnaissance studies and cooperated by having a por¬ 
tion of the special barrels prepared at Springfield Armory 
from steels supplied by contractors of Division 1 who had 
encountered difficulties in getting special barrels made prompt¬ 
ly by the machine gun barrel manufacturers. 


rel steels on the life and performance of stellite-lined 
barrels (with and without other features, such as 
plate ahead of the liner and special weight and con¬ 
tours) be pursued. 


24 5 3 Tests of Stellite-Lined Barrels 
Made of Special Steels 

Low-Alloy Steels 

The results of tests, extent of testing, and progress 
on preparation for test of barrels of low-alloy steels 
containing stellite liners 6 may be summarized as fol¬ 
lows : 

Templex Steel. Firing tests on four caliber .50 air¬ 
craft barrels of Templex steel (ASTM Specification 
A 193-44 T Grade B 14, a vanadium bearing SAE 
4140-type steel) with a 9-in. liner of Stellite No. 21 
showed that it is possible to fire a continuous burst of 
425 rounds without serious inaccuracy and with little 
warping or wilting of the barrel. On control tests with 
similar lined barrels of WD 4150 gun steel the barrels 
distorted and the barrel blew out in an attempt to 
fire such a long continuous burst. Thus use of the 
special steel gives a safety factor when emergency 
combat conditions necessitate a very long burst of 
fire. 

3% Molybdenum Steel. A firing test on a caliber .50 
aircraft barrel of a 3% molybdenum steel containing 
0.13% carbon with a 9-in. liner of Stellite No. 21 
showed that the accuracy-life and velocity-life ob¬ 
tained from this barrel were superior to those ob¬ 
tained with similar lined barrels of WD 4150 gun steel 
when fired on the same severe schedule (CGL-350 
schedule). 

2% Molybdenum Steel. Fifteen caliber .50 aircraft 
barrels were prepared for test from a steel having 
the composition: 2.05% Mo, 0.45% Mn, 0.10% Ni, 
0.03% Cr, 0.29% Si, 0.20% P, 0.015% Si, and 0.33% 
C. Five barrels had 9-in. liners of Stellite No. 21 in¬ 
serted. Two were fired and three were delivered to the 
Ordnance Department for its tests. The two barrels 
fired showed no improvement over WD 4150 steel be¬ 
cause the steel was not hardened. Five barrels with 
9-in. liners of Stellite No. 21 were induction hardened 
on both the outside diameter and on the bore and 
were delivered to the Ordnance Department for test. 


e All these stellite liners were of the design shown in Figure 1 
of Chapter 22. 


CONFIDENTIAL 





CHANGE OF BARREL STEEL TO ENHANCE PERFORMANCE 


483 


The hardening on the outside diameter was per¬ 
formed to test the strengthening effect of such a hard¬ 
ened layer in preventing barrel expansion under very 
severe firing conditions. Five barrels recessed to re¬ 
ceive 9-in. liners of Stellite No. 21 and induction 
hardened on the outside diameter and on the bore 
were chromium plated ahead of the liner recess. After 
this operation liners were inserted and the barrels de¬ 
livered to the Ordnance Department for test. The re¬ 
sults of the Ordnance Department tests of these several 
groups of barrels were not available when this sum¬ 
mary was written. 

Timken 1722A Steel. Springfield Armory prepared 
for Division 1 six 13-lb caliber .50 aircraft machine 
gun barrels from Timken 1722A steel, which is a low- 
alloy steel somewhat similar to ordinary WD 4150 gun 
steel but with superior high-temperature strength. 
Liners of special chromium-base alloys were inserted 
in four barrels (Section 17.4.3) and stellite liners in 
two barrels. All barrels were fired 5,000 rounds on the 
severe CGL-350 schedule with no failure. Firing 
should be continued. 

High-Alloy Steels 

The results of tests, extent of testing and progress 
on preparation for test barrels of high-alloy steels are 
summarized below. 

All of the high-alloy steels were difficult to machine 
and there is some doubt as to whether it would be 
practical to manufacture barrels for this reason. 
Another serious question is whether during war time 
sufficient critical alloying elements would be avail¬ 
able to permit the manufacture of a large number of 
barrels from such high-alloy steels. 

Chro-mow Steel Barrels. Two barrels of Chro-mow 
steel (5 Cr, 1.35 Mo, 1.25 W, 1 Si, 0.30 C) from the 
Crucible Steel Company, with 9-in. liners of Stellite 
No. 21 were subjected to firing tests. The results were 
somewhat inconclusive because the barrels had been 
bored oversize for chromium plating, but no plate 
was applied. After severe firing both the steel and 
liner showed good resistance to erosion but the barrel 
was inaccurate owing to the oversize bore. 

Peerless A Steel. Four barrels of Peerless A steel 
(9 W, 3.25 Cr, 0.25 V, 0.28 C) from the Crucible Steel 
Co., with 9-in. liners of Stellite No. 21 were tested. 
They had double the life of similar barrels of WD4150 
gun steel on the CGL-350 schedule. 

Silchrome XCR Steel. Silchrome XCR steel (24 Cr, 

5 Ni, 3 Mo, 0.45 C) from the Allegheny-Ludlum Steel 


Corp. has better high-temperature properties than 
gun steel and when tested as p, liner in the caliber .50 
heavy machine gun barrel showed better erosion re¬ 
sistance than WD 4150 gun steel. Thirty-two bars of 
this steel were sent to Springfield Armory by Divi¬ 
sion 1 for the preparation of special barrels. Twenty 
special caliber .50 aircraft barrels (some regular 10-lb 
and some special 13-lb barrels with 9-in. liners of 
Stellite No. 21) were to be prepared as well as six 
caliber .30 aircraft barrels with no liner. It was found 
impossible to drill or machine this steel because of 
hard spots in the steel supplied. 

Potomac Grade 81 Steel. Springfield Armory pre¬ 
pared three 13-lb caliber .50 aircraft barrels of 
Potomac Grade 81 steel (4.8 Cr, 1.28 W, 1.3 Mo, 
0.20 V, 0.32 C) from the Allegheny-Ludlum Steel Corp. 
These barrels, which contained 9-in. liners of Stellite 
No. 21 were sent to Crane Company for test. After 
termination of the NDRC contract at Crane, they 
were transferred to the Ordnance Department for 
test. 

HI A Grade 160 Steel. Springfield Armory pre¬ 
pared three 13-lb caliber .50 aircraft barrels of HI A 
Grade 160 steel (4.9 Cr, 4 Mo, 1.8 Ni, 0.44 C) from the 
Allegheny-Ludlum Steel Corporation. These barrels, 
which contained 9-in. liners of Stellite No. 21 were 
sent to Crane Company for test. After termination of 
the NDRC contract at Crane, they were delivered to 
the Ordnance Department for test. 

Grade 1081+ Steel. Springfield Armory attempted to 
prepare three 13-lb caliber .50 aircraft barrels from 
Grade 1084 steel (9 Cr, 1 Mo, 0.26 C) from the Alle¬ 
gheny-Ludlum Steel Corp. Two barrels were spoiled 
owing to the poor machinability of this steel. One 
barrel was completed and sent to Crane Co. for test. 
Owing to termination of the NDRC contract at 
Crane, this barrel was transferred to the Ordnance 
Department for test. 

Grade 1093 Steel. Crane Company obtained from 
the Allegheny-Ludlum Steel Corp. several bars of 
Grade 1093 steel. They intended to send them to 
Springfield Armory, which was to prepare 13-lb cal¬ 
iber .50 aircraft barrels. Owing to the termination of 
the NDRC contract at Crane Company, this steel 
was never sent to the Armory. 

XB Valve Steel. Geophysical Laboratory sent sev¬ 
eral bars of XB valve steel from the Allegheny-Lud¬ 
lum Steel Corp. to Springfield Armory. This steel has 
the composition 19.5 Cr, 1.35 Ni, 2.29 Si, 0.76 C. Two 
regular caliber .50 aircraft barrels with 9-in. liners of 
Stellite No. 21 were made and retained by the Ord- 


CONFIDENTIAL 




484 


BARRELS BOTH STELLITE-LINED AND CHROMIUM-PLATED 


nance Department after termination of the experi¬ 
mental program of Division 1. 

Exelloy Steel. Three regular 10-lb caliber .50 air¬ 
craft barrels of this steel were made for Crane Com¬ 
pany by Springfield Armory from Exelloy steel, 
which is a Crane Company steel containing 13% Cr, 
0.5% Ni, 0.11% C. One barrel without a liner gave a 
performance inferior to that of a regular WD 4150 
steel barrel. Regular Stellite No. 21 liners were to 
have been inserted in the other two barrels, but one 
barrel was spoiled in processing. A 6-in. external rein¬ 
forcing sleeve (Section 24.2) of Templex steel (see 
above) was fitted on the other barrel with a liner and 
it was subjected to a firing test. Its performance was 
comparable to a similarly reinforced WD 4150 steel 
barrel with a liner except that the Exelloy steel bar¬ 
rel ruptured, terminating the test. The resistance of 
this steel to impact from the bullets apparently is 
very poor. 

Potomac Hot Die Steel. Six regular 10-lb caliber .50 
aircraft barrels of Potomac Hot Die Steel (5 Cr, 1.7 
Mo, 1.3 W, 1 Si, 0.33 C) from the Allegheny-Ludlum 
Steel Corp. were made for Geophysical Laboratory 
by Springfield Armory. These barrels were for use in 
bore hardening prior to chromium plating but similar 
barrels could be used with liners. This steel is very 
similar to the Chro-mow steel already discussed. 

TK Hot Die Steel. Two regular 10-lb caliber .50 
aircraft barrels of TK Hot Die Steel (10 W, 3.8 Cr, 
0.22 Cb, 0.1 Mo, 0.06 V, 0.34 C) from the Carpenter 
Steel Co. were made for Geophysical Laboratory by 
Springfield Armory. The steel was first heat treated 
to about 35 Rockwell C. Both barrels were defective 
and were not fired. Pieces of this steel, which has good 
high-temperature properties, were used at Geophys¬ 
ical Laboratory to make the external reinforcing 
sleeves for regular WD 4150 steel barrels with stellite 
liners described in Section 24.2. 


24 6 OPTIMUM COMBINATION FOR BEST 
PERFORMANCE 

In summary we may say that to obtain the best 
caliber .50 aircraft machine gun barrel performance 1 
the following features should be combined: 

(1) Breech liner of Stellite No. 21, 9 in. long; 

(2) Choked-muzzle chromium plate ahead of the 
liner; 

(3) Proper change in weight and external contour 
of the barrel (total weight and distribution of metal 
consistent with cyclic rate of fire desired and max¬ 
imum permissible weight). 

(4) Barrel made from special heat-resisting steel 
with better elevated temperature properties than 
conventional machine gun barrel steel. 

(5) External or internal barrel cooling, a subject 
discussed in Section 5.7. The cooperation of the Air 
Forces working hand in hand with the developers of 
gun barrels, guns, and gun mounts will be necessary 
in order to combine effectively external barrel cooling 
with the other improvements. 

Attention is called here to some very suggestive 
firing tests (Section 17.4.3) of 13-lb caliber .50 aircraft 
barrels of a special steel containing liners of two dif¬ 
ferent chromium-base alloys. Because of their higher 
melting temperatures the chromium-base alloy liners, 
unlike those of stellite (Section 19.5.2), should per¬ 
form well in firings with higher velocity ammunition 
using double-base powder. Further development and 
tests of the chromium-base alloy liners should be 
made to evaluate their potentialities in machine gun 
barrels under hypervelocity conditions. 

f Melting or at least softening of the bullet jacket from fric¬ 
tional heat (Section 6.1.2) may impose an additional limitation 
on gun performance when the present gilding metal jacket is 
used. No material having greater strength at elevated temper¬ 
atures has been suggested that would not increase the frictional 
wear on the barrel. 


CONFIDENTIAL 






Chapter 25 

/ 

PILOT PLANTS FOR CHROMIUM-PLATING CALIBER .50 BARRELS 

By V. Wichum h and C. A. Marsh c 


251 DEVELOPMENT OF IMPROVED 
BARRELS 

ttihe rather rapid development of the improved 
-i- caliber .50 aircraft machine gun barrels described 
in Chapters 23 and 24 was made possible by the set¬ 
ting-up of pilot plants for small-scale production after 
other essentials of design had been determined by 
laboratory experiments. A brief chronological ac¬ 
count of the development which led to the adoption 
by Army Ordnance of these improved barrels to 
supersede the formerly standard steel barrel is deemed 
worth while in order to evaluate the success of the 
pilot-plant projects. 

The nitrided, chromium-plated caliber .50 aircraft 
barrel with choked muzzle was designed as a result of 
experiments to find a suitable hot-hard material to 
resist swaging and an adherent erosion-resistant plat¬ 
ing to prevent powder-gas erosion of the bore. 81 The 
choked muzzle was to increase the accuracy-life of 
the barrel. The plating and electropolishing proce¬ 
dures (described in Chapter 20) were worked out at 
the National Bureau of Standards in collaboration 
with the Geophysical Laboratory, Carnegie Institu¬ 
tion of Washington, where firing tests and examina¬ 
tion of fired barrels were carried out to serve as a 
basis for specifications, as described in Section 23.1.2. 

The techniques employed by the National Bureau 
of Standards were not necessarily practical for large- 
scale production. To prove the practicability of chro¬ 
mium-plating large quantities of caliber .50 barrels, a 
commercial plating company, the W. B. Jarvis Com- 
pany d was chosen in July 1944 to carry out pilot- 


a This chapter is based largely on an informal report sub¬ 
mitted on November 29, 1945, to the National Bureau of 
Standards by the Doehler-Jarvis Corporation and on the final 
report 142 from the Chrome Gage Corporation on Contract 
OEMsr-1444. The former report has not been distributed; and 
the latter has had a very limited distribution. 

b Engineer, Division 1, NDRC. (Present address: New 
York, N. Y.) 

c Geophysical Laboratory, Carnegie Institution of Washing¬ 
ton. (Present address: U. S. Geological Survey, Washington, 
D. C.) 

d Later, the Jarvis Division of the Doehler-Jarvis Cor¬ 
poration. 


plant operations resulting in limited production of 
barrels for testing purposes, under a subcontract with 
the Geophysical Laboratory. Necessary changes in 
the recommended procedure were made in order to 
conform to production methods. 

An extensive research program was carried out by 
the Geophysical Laboratory and the Jarvis Company 
to evaluate the importance of various steps in the 
plating procedure. On the basis of this research, final 
specifications and the procedure outlined in Section 
25.2.1 were established. 81 

Thereupon, several hundred barrels plated by the 
Doehler-Jarvis Corporation were delivered by Divi¬ 
sion 1 to the Army Ordnance Department for test. 
This barrel was finally adopted for Service use in 
January 1945. 294 

The need for immediately starting production of 
chromium-plated barrels led the Ordnance Depart¬ 
ment to request the assistance of Division 1 in guid¬ 
ing prospective Ordnance contractors. Thereupon, 
early in 1945 a contract was arranged with the 
Chrome Gage Corporation for an additional pilot 
plant to serve specifically as a control for Ordnance 
production contracts. 

During the setting-up of the latter pilot plant, the 
Doehler-Jarvis Corporation continued to produce 
barrels for test at the Geophysical Laboratory and 
for the Services. In April 1945, production was 
started under a separate Ordnance contract with this 
company. During the last two months of its operation, 
Doehler-Jarvis was chromium plating almost 700 
barrels daily on that contract. 

By the time that a small number of barrels had 
been produced by the Chrome Gage Corporation 
successful experiments with the “combination” bar¬ 
rel (Chapter 24) led the Ordnance Department to 
cancel orders in June 1945 for the nitrided, chromium- 
plated barrel with choked muzzle and to request 
production of the “combination” barrel. This barrel, 
which contained a 9-in. liner of Stellite No. 21 and 
which was chromium-plated ahead of the liner to 
give a choked-muzzle, was found to have both a long 
accuracy-life and a long velocity-life. In spite of the 
necessity for developing new plating techniques for 
the “combination” barrel, several hundred barrels 


CONFIDENTIAL 


485 



486 


PILOT PLANTS FOR CHROMIUM-PLATING BARRELS 


were plated both by Chrome Gage and Doehler-Jarvis 
before the surrender of Japan brought about the 
closing of both pilot plants. 

25.2 PILOT plant of the doehler- 
JARVIS CORPORATION 

25 2 1 Plating Procedure for Nitrided Barrels 

The chromium-plating procedure recommended by 
the National Bureau of Standards, as given in Sec¬ 
tion 20.2.3, is essentially the one followed by Doehler- 
Jarvis. Some changes Avere made as a result of firing 
tests on several lots of barrels that were plated, with 
variations in the procedures. 110 About 50 different 
experimental lots of barrels were plated, there being 
one to five barrels in a lot. The type of anode, the 
trivalent chromium content of the bath, times of plat¬ 
ing, current density, longitudinal distribution of chro¬ 
mium, length of reversal of current prior to plating, 
method of scrubbing the bore, and so on, were varied 
independently. The barrels were tested by firing and 
carefully examined. The resulting preferred proce¬ 
dure is listed below together with appropriate notes 
on the various steps. Also the changes that were made 
from time to time are noted. A diagram of the three 
principal steps in the process of producing the im¬ 
proved barrel from a standard barrel is shown as 
Figure 1 of Chapter 23. 

Decoppering was one of the steps listed in Section 
20.2.3. This was done, in the case of the nitrided bar¬ 
rels, before nitriding and so it is not included here. 
Also for a discussion of the nitriding and tests prior 
to plating, the reader is referred to Section 23.1.2. 

Outline of Procedure 

The steps in the procedure will not only be listed 
but will be grouped as follows: 

1. Receive, unpack, check quantity and type. 

2. Buff muzzle bearing to remove chromium 
plate. 

3. Number, giving lot number, Doehler-Jarvis 
identification “J,” and serial number, (Exam¬ 
ple—L2-J20). 

4. Paint barrel with stop-off lacquer. 

5. Clean barrel. 

a. Swab with solvent (Triad). 

b. Scrub with inhibited hydrochloric acid and 
pumice. 

c. Scrub with alkaline cleaner and pumice. 


d. Boroscope. 

e. Rinse and dry. 

f. Oil lightly. 

6. Gauge. 

a. Make out Record Sheet for each barrel and 
enter all information up to this point. 

b. Gauge lands and grooves at designated points 
and gauge centering cylinder. 

c. Enter above information on Record Sheet. 

d. Determine time for electropolishing and en¬ 
ter on Sheet. 

7. Clean barrel. 

a. Same as Step No. 5 except that boroscoping 
and final oiling is unnecessary. 

8. Rack for electropolish. 

a. Assemble fittings and anode, making sure 
fittings are clean, insulators are in good re¬ 
pair, and that anode is clean and straight. 
Use Saran tube over anode when inserting in 
barrel. This tube must be wiped with solvent 
each time before use. Anode must be checked 
for tension. 

b. Assemble hanger with above assembly, mak¬ 
ing sure that all contacting points are clean 
and that the barrel hangs straight. 

9. Electropolish (see Figure 1 of Chapter 23). 

a. Check temperature of solution (108-112 F) 
and see that agitation is constant and ade¬ 
quate. 

b. Place barrel in tank and make connections. 

c. Apply current, starting with 9-10 volts. Am¬ 
perage will drop after a short time. When 
this occurs, adjust current to 90 amp. 

d. Electropolish for designated time. Operator 
should place tag on hanger showing “Time 
In” and “Time Out.” Also a record of this 
operation is to be made on sheet provided for 
this purpose. 

10. Clean and unrack. 

a. Rinse with spent Cr 03 solution. 

b. Rinse with water. 

c. Unrack. 

d. Rinse. 

e. Scrub with pumice and cleaner. 

f. Rinse and dry. 

g. Oil lightly. 

11. Gauge. 

a. Gauge lands and grooves at designated 
points and record on Record Sheet. 

b. Calculate and enter amount removed on 
Sheet. 


CONFIDENTIAL 



PILOT PLANT OF THE DOEHLER-JARVIS CORPORATION 


487 


c. Calculate plating time and enter on Record 
Sheet. 

12. Clean. 

a. Same as Step No. 7. 

13. Rack for plating. 

a. Same procedure and precautions as under 
Step No. 8. 

14. Pretreat (anodic etch in chromic acid). 

a. Check temperature of pretreat tank (125 F) 
and check agitation. 

b. Place barrel in tank, make connections, and 
apply current for 5 min at 70 amp. 

15. Plate (See Figure 1 of Chapter 23). 

a. Check temperature of plating tank (122-123 
F) twice a shift and check agitation. 

b. Remove barrel from reversing bar (in pre¬ 
treat tank) and place on plating bar. Make 
connections and apply current at 70 amp for 
designated time. Tagging and recording as 
in Step No. 9. 

16. Clean and unrack. 

a. Same as Step No. 10 except that rinsing be¬ 
fore unracking is done with water only. First 
cold and then hot water is used. 

17. Gauge. 

a. Read lands and grooves at designated points 
and centering cylinder. 

b. Calculate and record thickness of plate. 

c. Calculate ratio for thickness at breech and 
muzzle and plating speed. 

d. Place all barrels, O.K. and rejects, in box or 
truck for removal to Final Inspection. 

e. Transfer Record Sheets to Final Inspection. 

18. Inspect finally. 

a. Boroscope. 

b. Check gauging records for completeness and 
compliance with specific dimensions and tol¬ 
erances. 

c Place O.K. barrels in space provided. 

d. Send rejected barrels back for stripping and 
replating or place with scrap barrels. 

e. Make proper disposition of all Record 
Sheets. 

f. Enter information on all Record Sheets on 
“Disposition” sheet. This sheet must show 
status of each barrel plated. 

19. Remove paint and prepare for shipment. 

a. List O.K. barrels according to lot and barrel 
number. 

b. Remove paint and thoroughly clean barrel 
inside and out. Clean threads and flutes at 


breech end with a wire brush to remove rust, 
scale, and chemical deposits. 

c. Paint an olive drab band 4 in. wide around 
each barrel' 2 in. behind front muzzle bear¬ 
ing. 

d. Dip barrel in rust-proofing oil, wrap in paper 
and pack. Check lot and serial number 
against list. 

Cleaning Operations 

Degreasing. During intervals when the barrels were 
not being treated in some one of the various solutions, 
it was necessary to keep them oiled to prevent rust¬ 
ing. The grease or oil had to be removed prior to some 
of the steps of the procedure. This was done by swab¬ 
bing the bore with Triad. 

Scrubbing. Before gauging the nitrided barrels prior 
to electropolishing it was necessary to scrub the bore 
to remove the brittle iron nitride layer. This opera¬ 
tion was performed with a tightly fitting patch or a 
wire brush on a cleaning rod, using powdered pumice 
and a dilute, inhibited hydrochloric acid solution; 50 
per cent concentrated HC1 by volume plus 10 ml per 
liter of “Rodine” inhibitor. A boroscope was used to 
differentiate between a well-scrubbed bore and one 
with a nitride layer. 

A scrub with alkaline cleaner was always employed 
after one with hydrochloric acid. Other scrubbing 
operations when acid was not used were usually car¬ 
ried out with the cleaner plus powdered pumice. The 
solution contained 8 oz/gal of a commercial cleaner. 

Rinsing. Thorough rinsing was always necessary 
after the barrels had been subjected to scrubbing, 
polishing, or plating solutions. After the electropol¬ 
ishing, the bores were rinsed with a small amount of 
spent chromic acid plating solution to avoid etching 
by sulfuric acid which occurs when the polishing solu¬ 
tion is diluted. After the electroplating, a hot rinse 
was used to facilitate the drying which had to be done 
as quickly as possible. A cold rinse was employed first 
to remove the bulk of the plating solution so the hot 
rinsing tank would not become contaminated. 

Oiling. As was stated above, oiling was necessary 
between some steps to prevent rusting. Finger print 
removing oil was applied lightly. 

Gauging Operations 

The gauging before electropolishing and electro¬ 
plating was carried out with a Sheffield Precisionaire 


CONFIDENTIAL 




488 


PILOT PLANTS FOR CHROMIUM-PLATING BARRELS 


gauge. This served not only as a routine check on the 
barrels during processing but enabled the operator to 
calculate the times the barrels should remain in the 
solutions in order to obtain dimensions that were 
within the specified tolerances. 

Preliminary gauging after chromium-plating to de¬ 
termine the diameter of the centering cylinder was 
done by means of plug gauges. The first of the final 
gauging operations made use of the Sheffield Preci- 
sionaire gauge. If the barrels were found to be within 
th£ dimensional tolerances, they were taken to the 
next operations (Nos. 18 and 19). If not satisfactory, 
they were regauged with a Federal Star Gauge. As a 
result of this step they were either accepted or re¬ 
jected or put into the “merit” classification. 

Electropolishing 

To remove sufficient metal from the bore to accom¬ 
modate the chromium plate that was to be added and 
to remove burrs and sharp corners, the bores of the 
barrels were electropolished using the solution de¬ 
scribed in Section 20.2.3. The solution, kept at 108- 
112 F, contained 50% by volume each of concentrated 
(96%) sulfuric acid and concentrated (75%) phos¬ 
phoric acid. Its specific gravity was maintained be¬ 
tween 1.68 and 1.76 by adding or evaporating water. 
A %rin. diameter copper rod plated Avith a lead-tin 
alloy (0.005 in. thick) was used as a cathode. The 
initial current applied was higher than the correct 
operating current (93 amp, Avhich gave a current den¬ 
sity of 250 amp per sq ft). If the initial current was 
low, followed by an increase, the bore might be etched 
instead of polished. Vigorous air agitation was neces¬ 
sary to prevent a taper. 

With the above procedure, the polishing rate was 
about 0.001 in. on the diameter in 10 min. The time 
usually required Avas slightly more than ten minutes. 

Anodic Etch Prior to Plating 

After the barrels were assembled in the plating 
bath a reversal of current prior to plating Avas applied 
for 5 min, a time determined by experimental work. 
This procedure, Avhich cleaned the bore surface and 
gave it a slight etch, Avas found to improve the adhe¬ 
sion of the chromium plate to the steel. 

Electroplating 

The barrels Avere plated with HC chromium (see 
Section 20.2.2) in such a manner that the deposit Avas 


thicker at the muzzle end than near the breech. This 
taper Avas automatically obtained AA-hen the proce¬ 
dure given beloAv Avas followed. The conditions AA'hich 
affect the amount of taper are discussed in Section 
20.2.3. 

The solution, kept at 122 ± 2 F, contained 250 ± 
10 g chromic acid (Cr0 3 ) and 2.5 g ± 0.1 anhydrous sul¬ 
furic acid per liter. During continuous use of the 
tanks, the amount of chromium reduced to the triva- 
lent state increased AAith a resultant substantial in¬ 
crease in the plating speed so that corrections in the 
plating time had to be made. Also, tapered anodes (% 
in. at the muzzle end and 3^8 in- at the breech end) 
Avere tried Avhen, as a result of operation, the amount 
of trivalent chromium in the bath affected the taper. 
The anode to be used Avhen the solutions were neAv 
Avas a %-in. steel rod plated with a lead-tin alloy 
(0.005 in. thick). Either a Y^-m. or %-m. anode aatis 
used depending on the age of the bath. The plating 
current Avas 69 amp, Avhich gave a current density of 
190 amp per sq ft. Air agitation insured a uniform 
temperature and bath composition. 

The assembling of the barrels, anodes, and fittings 
for plating Avas essentially the same^as for electropol¬ 
ishing. The barrels Avere plated with the muzzle end 
up and the current connection was made at the top of 
the anode. It Avas necessary to stop-off the first 3J4 
in. at the breech end of the coated steel anode AAdth 
Saran tubing so that the corresponding section of the 
chamber Avould not be plated. 

The plating rate for a neAv bath Avas about 0.001 in. 
per hour on the diameter at 3134 in- from the muzzle, 
Avhich is just ahead of the origin of rifling. 

The finished bore surface Avas to have a continuous 
bright plate that Avas free from nodules, blisters, flak¬ 
ing, or other surface imperfections. 

Fittings used to center and insulate the anodes and 
the current connections are shoA\m in Figures 3 and 4 
in Chapter 20. 


Simplified Procedure 

A simplified procedure, in Avhich several steps Avere 
omitted, Avas tried on one lot of barrels after the Ord¬ 
nance Department had approved some barrels plated 
by the above procedure. It was found to be satis¬ 
factory. 

The major changes Avere the omission of lacquer¬ 
ing (Step 4 of Section 25.2.1), and the omission 
of the scrubbing and gauging prior to electropolish¬ 
ing (Steps 5,6, and 7 of Section 25.2.1), since the latter 


CONFIDENTIAL 




PILOT PLANT OF THE DOEHLER-JARVIS CORPORATION 


489 


removes the nitride layer equally well. The barrels 
were merely degreased before being polished. The 
rinse with spent chromic acid solution after electro¬ 
polishing (Step 10a) was also omitted in the final pro¬ 
cedure. 


25 2 2 Dimensions of Nitrided, 

Plated Barrels with Choked 
Muzzles 

The diameter of the bore across the lands at both 
breech and muzzle ends of the regular, steel barrels 
from which the improved type of barrel was made by 
nitriding and chromium-plating was between 0.4990 
and 0.5010 in. In the latter type of barrel, however, a 
taper of bore diameter from the origin of rifling to the 
muzzle was found to be the feature essentially re¬ 
sponsible for the verjr great prolongment of accuracy- 
life. 

The bore dimensions after plating had to be held 
within specified limits for optimum performance. On 
the basis of research up to October 1944, which was 
prior to operations by Doehler-Jarvis, the dimen¬ 
sions 81 were specified to be as shown in the second 
column of Table 1. 


Table 1 . Final specifications for nitrided, chromium- 
plated caliber .50 aircraft barrels. 


Distance from 
muzzle (in.) 

Specified diameter (in.) 

Oct. 1944 June 1945 

m 

0.4950 + .0015 

0.4920 + .0040 

10 

0.4980 (max.) 

0.4950 + .0030 

15 

0.4985 (max.) 

0.4960 + .0030 

31H 

0.4985 + .0015 

0.4985 + .0025 

32.1 

0.512 + .002 

0.513 + .002 


The maximum value for the diameter at the muzzle 
was set at 0.4965 in. after it had been found that the 
muzzle had to be choked to at least this extent to ob¬ 
tain optimum performance. At the origin of rifling the 
diameter could not be less than 0.4985 in. because the 
excessive gas pressures which would develop when 
the bore in this region was smaller than this might 
cause primers to be blown. Between both ends of the 
rifled portion of the barrel a uniform taper was desir¬ 
able, but limiting dimensions were not set for posi¬ 
tions at equal distances throughout the bore. It was 
important, however, that there should not be a sharp 
taper right at the muzzle, hence maximum diameters 
were specified for positions at 10 and 15 in. from the 
muzzle end. 



RIGHT SECTION OF A CHROMIUM PLATED GUN BARREL 
1/2" FROM MUZZLE 



■- - ' 

RIGHT SECTION OF A CHROMIUM PLATED GUN BARREL 
10" FROM MUZZLE 




RIGHT SECTION OF A CHROMIUM PLATED GUN BARREL 
31" FROM MUZZLE 


Figure 1 . Varying thickness of chromium plate at dif¬ 
ferent positions along the bore of a nitrided and chrom¬ 
ium-plated caliber .50 barrel with choked muzzle. 
(From a Doehler-Jarvis report to Geophysical Lab.) 


CONFIDENTIAL 










490 


PILOT PLANTS FOR CHROMIUM-PLATING BARRELS 



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CONFIDENTIAL 


















































































































































































































































































































































































































































































































PILOT PLANT OF THE DOEHLER-JARVIS CORPORATION 


491 



Figure 3. Production unit in Chrome Gage pilot plant. (From OSRD Report No. 6517.) 


During the course of the work at Doehler-Jarvis, 
these specifications were modified slightly, on the 
basis of experience gained in the testing of plated bar¬ 
rels. The final specifications as of June 6, 1945, the 
date the orders for these barrels were canceled be¬ 
cause of the development of the stellite-lined, chro¬ 
mium-plated barrel, described in Chapter 24, are given 
in the third column of Table 1. A comparison of this 
column with the second one shows the greatest change 
in specified diameters was at the muzzle. It was found 
that better barrel performance was favored with 
smaller diameters than 0.4950 in. This dimension 
could be even less than 0.4920 in. but in this case 
plating time would be too much longer than was justi¬ 
fied by the increase in barrel performance. 

The taper was obtained by chromium-plating the 
bore under carefully controlled conditions so that the 
deposit was thicker at the muzzle than at the breech. 
Figure 1 shows the variation in thickness from muzzle 
to breech in a typical barrel. The minimum thickness 
was the only one specified; thus, on any land at the 
origin of rifling it was to be 0.0020 ± 0.0005 in. 


The large amount of heat input (Section 13.2.5) to the 
steel in the cylinder of the caliber .50 machine gun 
barrel causes a decrease in diameter of about 0.002 
in. When the cylinder becomes constricted, stoppages 
may occur as a result of the failure of the cartridges 
to seat. In a regular steel barrel, constriction is not a 
problem because the unprotected steel is subject to 
erosion by the powder gases when the barrel is first 
fired, but in chromium-plated barrels the cylinder 
has to be made larger. The stock barrels, therefore, 
were reamed out before nitriding and plating. As 
can be seen from Table 1, the diameter (32.1 in. from 
muzzle) specified in October 1944 was not large 
enough. 

25 2 3 "Combination” Barrels 

Changes had to be made in the procedure described 
above for nitrided barrels when caliber .50 aircraft 
barrels that were to contain a stellite liner were 
chromium-plated. Only the portion ahead of the liner 
was to be plated. These barrels were not nitrided. 


CONFIDENTIAL 















492 


PILOT PLANTS FOR CHROMIUM-PLATING BARRELS 


If they had been proof-fired, it was necessary to de¬ 
copper them before the standard plating procedure 
was applied. Not enough work was done before the 
termination of the program to determine the best 
production method of plating these barrels. The 
difficulties encountered are discussed in Section 
20.2.3. 

25,2 4 Accomplishments of the 

Doehler-Jarvis Pilot Plant 

The operation of the Doehler-Jarvis pilot plant 
made at least three important contributions to the 
project. 

1. It provided the several hundred barrels necessary 
for testing by the interested Army and Navy Services, 
prior to acceptance of such barrels for production. 

2. It made possible the testing on an adequate 
scale of the large number of variables involved in pre¬ 
paring and plating the barrels, thus permitting choice 
of the most effective methods. 

3. It demonstrated that a well-managed and effi¬ 
ciently operated plant could carry out on a production 
scale the relatively complex and precise plating opera¬ 
tions involved, with uniformly good results, a very 
low percentage of rejections, and at reasonable costs. 


?53 PILOT PLANT OF THE CHROME 
GAGE CORPORATION 

25,31 Description of Project 

When Army Ordnance had outlined a program for 
the large-scale production of improved caliber .50 
aircraft machine gun barrels, the Chrome Gage Cor¬ 
poration of Philadelphia was chosen as the commer¬ 
cial plant to confirm or improve the plating technique 
developed at the pilot plant of the Doehler-Jarvis Cor¬ 
poration, and to operate a commercial plant. 

Under the terms of the contract 10,000 caliber .50 
aircraft machine gun barrels were to be nitrided and 
chromium-plated, the nitriding to be done by a sub¬ 
contractor. A pilot plant was to be installed to develop 
methods which would be better suited to quantity 
production than the existing ones, would yield more 
uniform results, and would substantially reduce the 
cost of plating caliber .50 gun barrels with hard 
chromium. The following program was set up. 

1. Design and equip an experimental plant or unit 
having a capacity of plating twenty barrels simul¬ 
taneously under varying conditions. 

2. Design and equip a pilot production plant or 
unit having a capacity of plating 100 barrels simul¬ 
taneously. 



Figure 4. Receiving, shipping, and final inspection areas of Chrome Gage pilot plant. (From OSRD Report No. G517.) 

CONFIDENTIAL 









PILOT PLANT OF THE CHROME GAGE CORPORATION 


493 



Figure 5. Electropolishing tank of production unit in Chrome Gage pilot plant. (From OSRD Report No. 6517.) 


3. Develop testing methods, inspection procedure, 
and specifications for all stages of processing from the 
time barrels were received until they were test fired. 

4. Explore possible improvements in testing and 
processing through research and development, and 
try same in the experimental unit before adoption, in 
production manufacture. 

5. Evolve methods of handling in production with 
the least expense of labor. 

6. Provide methods of checking and inspection, 
which would prevent faulty processing and furnish 
the necessary information by which results might be 
analyzed. 

7. Carry on experimental plating operations. 

The net result of this program was to be a check on 

the practicality of the tentative specifications 292 issued 
by the Ordnance Department for nitrided and chro¬ 
mium-plated caliber .50 aircraft machine gun barrels 
with choked muzzle (Chapter 23). Those specifica¬ 


tions had been based on the experience of theDoehler- 
Jarvis Corporation. Before production could get 
under way, the scope of the project was expanded to 
include the plating of the chromium-plated caliber 
.50 barrels with stellite liners at the breech end 
(Chapter 24), also in accordance with tentative Ord¬ 
nance Department specifications. 

25,3,2 Plant Installation 

When a termination order was received on Au¬ 
gust 17,1945, the production line of the pilot plant was 
not quite completed. It had been designed to handle 
500 barrels a day. An old dye house had been exten¬ 
sively remodeled in accordance with the floor plan 
shown in Figure 2. Two views of the partially com¬ 
pleted installations are shown in Figures 3 and 4. The 
electropolishing unit, which was complete except for 
some minor electrical work, is shown in Figure 5. 


CONFIDENTIAL 






494 


PILOT PLANTS FOR CHROMIUM-PLATING BARRELS 



25 ' 0 ' 


Figure 6. Floor plan of experimental plating unit in Chrome Gage pilot plant. (From OSRD Report No. 6517.) 


An experimental plating unit suitable for a large 
range of work had been in operation slightly over 3 
months. Its floor plan is shown in Figure 6 and an in¬ 
terior view of it in Figure 7. 

Noteworthy among the designs of equipment were 
ones for the following: a plating fixture and clamp for 
the gun barrel to provide for concentricity of anode, 
good electrical contact, and ease of assembly and 
handling; an efficient pneumatic pumicing unit for 
scrubbing the bores; a horoscope inspection unit; an 
improved star gauge; and an electrical bore gauge 
involving a new application of a strain gauge, as 
is described in Section 25.3.5. Groups of barrels 
were handled by the overhead conveyor shown in 
Figure 8. 

Interpretation of test results was facilitated by the 
recording methods devised. Samples of Form CGC 


No. 1, “Barrel Inspection Before Plating” and of 
Form CGC No. 2, “Gauging Record” are shown in 
Figures 9 and 10, respectively, with appropriate data 
filled in for actual barrels. These two forms were to be 
used for all barrels put through the production line, 
as well as for experimental barrels. 

In addition a “Graphic Record” was maintained 
on Form CGC No. 3 (Figure 11) for each barrel in the 
experimental plating unit to show the bore diameter 
through the length of the barrel initially, after electro¬ 
polishing, after plating, and after a firing test. The 
conditions for electropolishing and electroplating were 
recorded on the same form. Two other forms were 
also used for the experimental barrels to record the 
details of the firing test and the appearance of the 
bore surface as shown by a horoscope examination 
after the test. 


CONFIDENTIAL 











































































































PILOT PLANT OF THE CHROME GAGE CORPORATION 


495 


25 3 3 Barrels Handled and Processed 

Nearly 1,300 barrels were handled in one or more 
stages of the plating process, either in the experi¬ 
mental unit or in preliminary trials of the production 
line. 

Of these barrels, which had been nitrided and were 
to be chromium-plated, 382 were rejected before elec- 
tropolishing, 330 were rejected after electropolishing, 
and 266 after plating. Finally, 308 barrels were ac¬ 
cepted, based on dimensions and appearance. Some 
barrels had also been rejected before nitriding for the 
following reasons: (1) did not pass straightness test, 
(2) had rough or torn bullet seats, and (3) showed 
poor machining in rifling. Many doubtful barrels 
however, were considered acceptable, which accounts 
for the large number of rejections in further process¬ 
ing. Those rejected before electropolishing showed 
defects which it was not possible to remove by the 


scrubbing or cleaning operation. Similarly, most of 
those showing machining defects which had not been 
removed by either scrubbing or electropolishing were 
rejected, while a small number failed to meet dimen¬ 
sional requirements. 

Those rejected after plating fell into four categories, 
namely (1) failed dimensionally, (2) showed stains 
and miscellaneous irregularities, (3) had small pits, 
and (4) had lands rippled by machining marks. 

About 100 barrels with a liner recess were plated. 
Most of this plating, however, was carried out in or¬ 
der to develop a technique for the production line. 
Only 20 of these barrels were processed according to 
production methods. Eighteen were considered to be 
satisfactorily plated. 

25,3,4 Plating Procedure 

The procedure for chromium-plating the nitrided 



Figure 7. Experimental plating unit in Chrome Gage pilot plant, showing electropolishing and strip tanks in the left 
foreground, scrubbing table in Center foreground, rinsing tank in right foreground, and plating tank in background. 
(From OSRD Report No. 6517.) 


CONFIDENTIAL 





















496 


PILOT PLANTS FOR CHROMIUM-PLATING BARRELS 



Figure 8. Caliber .50 gun barrels suspended from overhead conveyor in Chrome Gage pilot plant, with rinsing tanks in 
foreground and plating tank in background. (From OSRD Report No. 6517.) 


barrels was essentially the same as that used by 
Doehler-Jarvis, involving the steps given in Section 
25.2.1. As had been done at the Doehler-Jarvis pilot 
plant, an effort was made in the experimental unit of 
the Chrome Gage plant to determine ways in which 
the procedure might be altered in order to increase 
production economically while maintaining or in¬ 
creasing the good performance of the barrels pro¬ 
duced. 

A rather extensive program was carried out to de¬ 
termine the feasibility of plating a group of barrels 
with only one current control device. In the original 
setup a separate one was needed for each barrel. 
The conclusions drawn from these experiments were 
that equal division of current could not be expected 
unless the surface resistance of each barrel was iden¬ 
tical, which could not be achieved without a rigid 


control over the cleaning and degreasing of the 
barrels. 

In the processing of some barrels, the scrubbing 
operations prior to plating were omitted, and in the 
case of others the electropolishing was omitted. In 
the absence of firing tests of those barrels no conclu¬ 
sions could be drawn. 

In plating a barrel that was to receive a stellite 
liner, a dummy steel liner usually was inserted in the 
liner recess according to the method used by Doehler- 
Jarvis, which is described in Section 20.2.3. Experi¬ 
ments were performed in order to improve the plating 
at the junction of the liner and the steel bore, but no 
improvement resulted. Finally plating was carried 
out without the dummy liner. The anode was stopped 
off opposite the portion of the barrel that was not to 
be plated. These barrels were not test fired, but there 


CONFIDENTIAL 




















PILOT PLANT OF THE CHROME GAGE CORPORATION 


497 



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CONFIDENTIAL 






































































































498 


PILOT PLANTS FOR CHROMIUM-PLATING BARRELS 


BREECH 
36" 31.25' 


POSITION ALONG AXIS OF BARREL 
25" 20" 15" 10" 


MUZZLE 
5" 1.5" 0 



Figure 11. CGC Form No. 3, “Graphic Record/’ showing bore diameter of caliber .50 barrel initially, after electro¬ 
polishing, after plating, and after having been test fired. 


were definite indications that it would be possible to 
plate barrels without dummy liners and obtain satis¬ 
factory plating results. 

In all these experiments on the barrels to contain 
liners, the barrels were electropolished with the breech 
end up so that the slight taper produced would be in 
the same direction as the taper produced later by the 
plating operation and thus the amount of plating 
necessary to offset a reverse taper, produced by pol¬ 
ishing with the muzzle end up, would not be required. 

Current Distribution Along a Cylindrical 
Anode 

A formula for the distribution of the current along 
the anode, assuming a uniform resistance of both 
anode and solution, was developed. Equations (1), 


(2), and (3) express the current and voltage across 
the solution and the current through a section of the 
anode, in terms of the geometry shown in Figure 12. 

7 T sinh a(l — x) m 

lx = 7o_ ^kThal“* (1) 

. _ al 0 Cosh a (l — x) 

. lx ird sinh al 

In these equations I is the current, i the current den¬ 
sity, E the voltage across the solution, E P the voltage 
of polarization, l the leagth (in cm) of exposed anode, 
d diameter of bore, a 2 the ratio r/p (where r is the 
resistance per unit length of the anode and p the 


CONFIDENTIAL 
























































RECOMMENDATIONS 


499 


resistance of a ring of solution 1 cm long), and x the 
distance in centimeters along the tube measured 
from the end where the current enters. However, the 
experience in plating caliber .50 barrels at the Na¬ 
tional Bureau of Standards, as described in Chapter 
20, casts doubt on the validity of the assumption of 
uniform resistance of the solution. Perhaps in “pump 
plating” the solution resistance may be sufficiently 


- //h r 



Figure 12. Geometrical relations in a gun barrel being 
plated with a cylindrical anode. (From OSRD Report 
No. 6517.) 

uniform for the analysis to apply. In plating without 
circulation of the solution, differences in temperature 
between the bottom and top of the barrel cause 
changes in both the solution resistance and cathode 
efficiency that tend to compensate each other as far 
as their effects on the current density are concerned, 
but the extent of compensation in any particular case 
is not predictable. 

25,3,5 Electrical Bore Gauge 

An electrical bore gauge, which was a type of strain 
gauge, was developed to permit the gauging of rifled 
and tapered bores with greater accuracy than is pos¬ 
sible with other types of gauges. The range of the in¬ 
strument was great enough so that it could be used to 
gauge eroded barrels as well as new ones. 



Figure 13. Top view of strain gauge beam removed 
from head of electrical bore gauge (magnified 4X). (From 
OSRD Report No. 6517.) 

This gauge makes use of a cantilever beam, shown 
in Figure 13, containing a sensitive electrical strain 
gauge that records the amount of motion of a feeler 
point on its end. This feeler point makes contact with 


another feeler supported by a spring fastened to the 
cover. The spring is three-pronged, the ends of the 
prongs on either side being curved so as to engage the 
grooves of the rifling and to center the feeler point on 
the land. On either side of the gauge head body is a 
fixed contact point. 

The gauge head is mounted in the end of a tube 
mounted on ball-bearings on a vertically adjustable 
support, which slides on a base on which are also 
mounted the vertically adjustable gun barrel rest, as 
shown in Figure 14. After calibration the gauge head 





Figure 14. Head of electrical bore gauge in position for 

insertion in a barrel. (From OSRD Report No. 6517.) 

is moved inside the bore, following the rifling. The 
gauge is connected to a potentiometric voltage indi¬ 
cator, on the dial of which is given a practically in¬ 
stantaneous measurement of the bore diameter at 
any point in the bore. The 30-in. scale is uniformly 
graduated in one thousandths of an inch, and each 
thousandth has 20 subdivisions, each representing 
0.00005 in. Thus the magnification is approximately 
1,250 to 1, for a gauging range of 0.025 in. Over this 
entire range the gauge will duplicate readings to 
within 0.0001 in. 

It was planned eventually to use the electrical bore 
gauge in connection with a potentiometric recorder. 
Then a continuous written record of the diameter of 
the bore could be obtained in a few seconds for the 
whole length of a caliber .50 barrel from the centering 
cylinder to the muzzle. 

25 4 RECOMMENDATIONS 

Future operation of a pilot plant for the chromium 
plating of machine gun barrels calls for four principal 
activities, as follows: 

1. Experiments to develop a technique for plating 


CONFIDENTIAL 

















500 


PILOT PLANTS FOR CHROMIUM-PLATING BARRELS 


“combination” barrels with stellite or other erosion 
resistant liner already inserted; 

2. Experiments looking toward simplification of 
the plating procedure; 


3. Development of a reasonably reliable gauge for 
measuring the thickness of the plate, in order to solve 
the problem of checking the concentricity of the plate 
with the bore. 


CONFIDENTIAL 




PART VII 

HYPERVELOCITY GUNS AND PROJECTILES 



When men are arrived at the goal, they should not turn 
back. —Plutarch 

“Of the Training of Children” 


CONFIDENTIAL 








Chapter 26 

i 

SHORT LINERS AND OTHER DESIGN FEATURES OF^ GUN TUBES 

By William H. Shallenberger a 


261 INTRODUCTION 

E rosion-resistant materials have been devel¬ 
oped to increase the life of gun barrels, ’as de¬ 
scribed in Chapter 16. For various reasons, however, 
it is not practical to construct complete barrels of 
such materials. Therefore gun barrels and tubes still 
need to be made of steel into which short erosion-re¬ 
sistant liners are inserted at the breech end. In gen¬ 
eral, the insertion of such liners introduces problems 
not encountered in tubes of monobloc construction or 
in ones containing full-length steel liners. (Figure 1.) 

The improvements made in machine gun barrels 
by the use of stellite liners have been discussed in 
Chapter 22. One of the purposes of the present chap¬ 
ter is to describe various methods that have been 
used to insert stellite liners in machine gun barrels as 
illustrations of how the general methods in use for 
designing gun tubes need to be modified to utilize 
erosion-resistant materials. The experiments already 
conducted by Division 1 with replaceable steel liners 
(Section 26.3) point the way toward one means of 
utilizing erosion-resistant materials in medium cali¬ 
ber guns. 

The specific designs suggested for a molybdenum 
liner for a 3-in./70-cal. gun are described in Section 
33.1.3. One of the special features of them is the 
stave-type of construction, already described in Sec¬ 
tion 18.5.2. An analysis of the stresses in these liners 
is presented in Section 26.4, following a review of the 
usual formulas for the stresses in shrunk-in liners. 
Consideration is also given to the effect on liner inser¬ 
tion of a difference between the coefficient of thermal 
expansion of steel and the liner material, such as 
molybdenum. 

The stress distribution in liners may also be in¬ 
fluenced by the design of rifling and the use of 


a Engineer, Engineering and Transition Office, NDRC. 
(Present address: Granada Hill, San Fernando, California.) 
While this chapter was being edited, which was after Mr. 
Shallenberger had left NDRC, several small additions were 
made to it, in order to take care of related subjects not treat¬ 
ed elsewhere in the volume: Section 26.3 was expanded by Dr. 
J. F. Schairer; Sections 26.5.1 and 26.5.2 were added by the 
Editor; and Section 26.6.3 was prepared by MissC. A. Marsh. 
(Editor's note.) 


pre-engraved projectiles, as discussed in Section 
26.6. 

26 2 MEANS OF HOLDING RESISTANT 
LINERS IN PLACE 

26-21 General Resume 

When a liner is inserted into a gun barrel it is nec¬ 
essary that the liner be held rigidly in place. The ac¬ 
tion of the projectile, in being accelerated down the 
bore, applies forces on the liner tending to move it 
forward and to rotate it. If the liner moves forward, 
the liner and the barrel at the forward joint become 
upset and constrict the bore, while rotation of the 
liner moves the rifling out of alignment and causes 
double-engraving of the projectile. 



Figure 1. A successful gun liner must follow the 
momentary changes in dimensions of the whole gun 
tube caused by passage of the projectile without under¬ 
going permanent set and without cracking. 

Various methods are used to secure the liner in the 
tube. Two of these methods produce an interference 
fit between the tube and liner, so that a high radial 
pressure exists at the interface. In the case of ma¬ 
chine gun barrels movement has been prevented me¬ 
chanically. Also it has been proposed to hold the liner 
by bonding it to the tube. The insertion of erosion- 
resistant liners for test purposes has already been de¬ 
scribed in Section 11.2.2. 

26 2 2 Shrink-Fit of Tube and Liner 

The means developed by the Crane Company for 
Division 1 for inserting and securing stellite liners in 
caliber .50 machine gun barrels was subsequently 


CONFIDENTIAL 


503 























504 


SHORT LINERS AND OTHER DESIGN FEATURES OF GUN TUBES 


used for nearly all production of these barrels during 
World War II. This method, as described in Section 
22.2.1, consisted essentially of boring out a recess at 
the breech end of the barrel extending approximately 
8 in. forward of the origin of rifling, and shrink-fitting 
a rifled cylindrical liner into the recess. A shoulder 
near the breech end of the liner butted against a cor¬ 
responding shoulder in the barrel to prevent forward 
motion of the liner. Figure 1 of Chapter 22 shows the 
assembly of this type of caliber .50 liner and barrel. 

The Savage Arms Corporation, on a production 
contract with the Ordnance Department, employed 
an assembly method similar to that of the Crane 
Company, with a few exceptions. Rifling of the barrel 
was done after assembly, using the rifled liner as a 
guide for the rifling tool. When it was decided to 
chromium-plate the barrel ahead of the liner (Chapter 
24) it became necessary to rifle the barrel before plat¬ 
ing and assembly, and use an indexing arbor during 
assembly to align the rifling. Instead of tack-welding 
the retainer to the barrel to prevent unscrewing, the 
threads of the retainer were silver-soldered to the 
threads in the barrel. 

26 2 3 Expansion of the Liner by Draw 
Rifling 

Interference between the barrel and liner was 
used 135 to prevent movement of the liner, but this 
method produced the interference after assembly 
rather than before. To obtain the desired interfer¬ 
ence, the liner was expanded sufficiently to give it a 
permanent set by forcing a plug through it. The ex¬ 
terior of this plug was so shaped that at the same 
time it expanded the barrel, it also formed the rifling. 
Two types of liner assemblies were prepared in this 
way by Remington Arms for the caliber .30 Brown¬ 
ing machine gun, M2-AC, M1917A1, and M1919A6. b 
The first was copied from the Crane design for caliber 
.30 (Section 22.5.1) except that the rifling and cham¬ 
bering operations were not performed prior to assem¬ 
bly, nor were interference fits used. After assembly 
the threaded breech end was screwed into a face plate 
on a La Pointe hydraulic push broach. The tungsten- 
carbide rifling plug or “button,” pushed through the 
barrel with a rotary motion, required about 5 sec to 
pass through the barrel. To reduce friction and pre¬ 
vent galling during rifling, the bore surface was given 

b Experimental work along the same lines was also done on 
caliber .50 machine gun barrels but did not progress as far as 
that on the caliber .30. 136 


a light plating of copper or indium. This plating was, 
of course, removed after rifling. 

Helical grooves on the outside of the plug formed 
the desired rifling by a swaging action. Displacement 
of metal caused the liner and retainer to take a per¬ 
manent expansion, thus locking them securely in the 
barrel. Since the amount of expansion and elastic re¬ 
covery are dependent upon the wall thickness, it was 
necessary that the barrel contour be cylindrical be¬ 
fore rifling, to prevent a taper in the finished bore. 
The outside of the barrel was machined to the de¬ 
sired contour after rifling. 

It was felt that the high interface pressure result¬ 
ing from swage-rifling would be sufficient to prevent 
forward motion of the liner and that the shoulder 
could be eliminated. Therefore, an alternate design 
was made as shown in Figure 2. Cylindrical bar stock 
was bored through with a diameter of 0.300 in. and 
then counterbored at one end to a diameter of 0.430 
in. for the liner. The liner and a steel ring were also 
bored through and turned to an outside diameter of 
0.430 in. After inserting the liner and retainer ring 
and plating the inside surface, the assembly was 
swage-rifled, as described previously. The chamber 
was then reamed in the normal manner. The steel re¬ 
tainer ring could have been eliminated, but was used 
to reduce the amount of stellite required and to pre¬ 
vent the stellite from coming to a feather edge, 
which might break off easily. 

This method of assembly is very inexpensive on 
account of the small amount of machining required, 
and also effects considerable savings in stellite by 
eliminating the flange at the end. One serious dis¬ 
advantage of swage-rifling the entire assembly is that 
it does not permit chromium-plating the barrel before 
assembly, because the plate would not withstand the 
rifling process. As described in Chapter 24, such plate 
ahead of the liner more than doubles the useful life of 
a stellite-lined caliber .50 machine gun barrel. Its 
value for the caliber .30 barrels was not determined 
by Division 1. Because of the reduced severity of 
erosion in this gun, compared with the caliber .50, it 
is questionable whether chromium plate ahead of the 
stellite liner is necessary. 

26 2 4 Mechanical Retention of the Liner 

The methods of inserting and retaining stellite 
liners in caliber .30 machine gun barrels previously 
discussed relied upon interference between the liner 
and the barrel. Mechanical methods of holding the 


CONFIDENTIAL 





MEANS OF HOLDING RESISTANT LINERS IN PLACE 


505 


LINER BUSHING 


LINER 

= 


BARREL 


DRAW RIFLED BARREL ASSEMBLY 

..... 


c 


Figure 2. Caliber .30 Browning machine gun barrel assembly, M1917A1-R1, with straight liner of Stellite No. 21; 
Remington design. (This figure was taken from NDRC Report No. A-463.) 


FINISHED BARREL 



liner were also developed. 127 By avoiding shrink-fits, 
it was felt that larger machining tolerances could be 
used, thus simplifying and speeding production. 

One promising design, shown in Figure 3, consisted 
of counterboring the breech end of the barrel to the 
depth necessary to accommodate the liner. A flange¬ 
less liner was inserted in the recess and retained by a 
simple set screw. Rifling of the liner and barrel could 
be done separately before assembly, or the liner could 
be rifled before assembly and used to guide the rifling 
bar for rifling the barrel. Although this method per¬ 


formed satisfactorily when using a steel liner, it was 
not tried with a stellite liner, because of the possibil¬ 
ity of crumbling of the feather edge at the breech end 
of the stellite. Subsequently, the Kelsey-Hayes 
Wheel Company, on a production contract for caliber 
.50 barrels, adopted a design using such a feather- 
edge without difficulty. Hence the design presumably 
would be practical for stellite. 

To avoid this feather-edge the design shown in 
Figure 4 was devised. In this design, the barrel and 
liner were pre-rifled and the liner inserted with a loose 



Figure 3. Caliber .30 Browning machine gun barrel assembly with flangeless liner; Johnson design. (This figure has 
appeared as Figure 2 in NDRC Report No. A-455.) 


CONFIDENTIAL 








































506 


SHORT LINERS AND OTHER DESIGN FEATURES OF GUN TUBES 



Figure 4. Caliber .30 Browning, machine gun barrel assembly, M2, with stellite liner staked with pins; Johnson de¬ 
sign. (This figure has appeared as Figure 4 in NDRC Report No. A-455.) 


or light press-fit. To determine what tolerances were 
required, several assemblies were made with different 
clearances over a range of several thousandths of an 
inch. All performed satisfactorily, indicating that 
close tolerances were not required. Forward motion 
of the liner was prevented by a shoulder on the liner, 
which seated on a corresponding shoulder in the bar¬ 
rel. To prevent rotation of the liner, the breech end of 
the liner was faced at an angle and the forward end of 
the chamber piece was faced at a corresponding- 
angle. The chamber piece was held in place and pre¬ 
vented from rotating by two tapered pins that se¬ 
cured it to the barrel. 

26 2 5 Bonding of Liner to Barrel 

Various methods of bonding a stellite liner to a 
machine gun barrel were proposed. Samples of stellite 
copper-brazed to gun steel prepared by salt-bath 
brazing showed satisfactory bond strength. 0 There¬ 
fore designs were made and a program formulated for 
the experimental production of stellite-lined caliber 
.50 barrels assembled by brazing. Two types were to 
be built and tested. One type was to be brazed over 
the entire length of the liner and separate chamber 
piece. In the other type, the stellite was to be bonded 
(by casting) to a steel end and this steel end to be 
brazed to the barrel. After assembly, the liner and 
barrel were to be swage-rifled by the process de¬ 
scribed in Section 2G.2.3. It is recommended that this 
unfinished project be completed. 

In other experiments stellite was bonded to steel 
by the process commonly known as the Al-Fin proc¬ 


c This work was carried out by the A. F. Holden Co. under 
Contract OEMsr-1473. Because the work under that contract 
did not progress much beyond the planning stage, the final 
report contained few details. Therefore it has not been issued 
as a formal NDRC report. 


ess. d This process has been used very successfully to 
bond aluminum muffs or finned surfaces to steel cyl¬ 
inders for aircraft engines. 559 However, tests showed 
serious weakening of the joint at high temperatures, 
in the case of the stellite-steel assembly. 

Stellite has been successfully bonded to steel by 
centrifugal casting into steel tubes. 125 These tubes 
with their stellite linings were experimentally assem¬ 
bled into caliber .50 barrels by standard insertion 
methods. The firing tests 80 showed that a stellite lin¬ 
ing prepared in this way gave a performance equiv¬ 
alent to that of an investment-cast liner. Attempts 
were to have been made to cast stellite centrifugally 
directly into barrel forgings bored out to receive it, 
either as short breech linings or as full-length linings. 
On account of termination of Division l’s experimen¬ 
tal program in 1945, this project was not completed. 
It is recommended that the Services undertake to 
finish it. 

26 3 INSERTION OF A REPLACEABLE 
STEEL LINER 

26,3,1 Purpose of a Replaceable Steel Liner 

The erosion of steel gun tubes as described in 
Chapter 10 is particularly severe at and near the ori¬ 
gin of rifling; but significant erosion also occurs for 
several calibers ahead of this region. This enlarge¬ 
ment of the bore by causing a rapid and undesirable 
drop in muzzle velocity and in some cases malfunc¬ 
tioning of the ammunition limits the useful life of the 
gun tube. 


d This work was carried out by the Al-Fin Corporation 
under Contract OEMsr-1494. Because the work under that 
contract did not progress much beyond the planning stage, the 
final report contained few details. Therefore it has not been 
issued as a formal NDRC report. 


CONFIDENTIAL 
































INSERTION OF A REPLACEABLE STEEL LINER 


507 


In 1942, before the erosion-resistant metals and 
alloys described in Chapter 16 had been developed, it 
was felt that the life of steel gun tubes, particularly 
under severe conditions of firing, might be extended 
considerably by the use of a series of short, inter¬ 
changeable, replaceable, steel breech-liners. The Ger¬ 
man 88-mm antitank gun was so constructed that its 
liner could be removed in three sections. 225 ’ 246 The 
development of a successful design and tests of liners 
of this type are described in the next two sections of 
this chapter. A modification of such a replaceable 
steel liner was used as a means of testing erosion- 
resistant materials in cannon tubes. For this purpose 
the erosion-resistant material may be applied as a 
bore lining or liner within the steel liner. Stellite was 
thus tested in a 37-mm gun, as described in Section 
33.1.2. 

26 3 2 Development of the Liner Design 

For the initial experimental work on the design of a 
replaceable steel liner, the 75-mm gun was chosen 
because of its ready availability. Three types of lin¬ 
ers, 10 calibers long and about 0.5-in. average thick¬ 
ness, were designed and constructed at Crane Com¬ 
pany, 80 and were tested at Aberdeen Proving 
Ground. 231 

The first two types were elliptical in cross section 
to prevent rotation. This involved difficult and ex¬ 
pensive machining operations. The first type failed 
by permanent expansion at the forward end, which 
was too thin, and by moving backward slightly upon 
firing. A ring of “copper” formed at the forward 
joint, and the liner could not be replaced until the 
copper had been machined out. The second type 
failed by radial expansion of the chamber section 
during the firing of five rounds. Removal of the liner 
in the field was impossible. 

The third and successful type of liner was of circu¬ 
lar cross section over the entire length, and had an 
external taper except for a thick flange at the breech 
end, which had the same outside diameter as the 
tube. The liner was secured to the tube by socket- 
head cap screws set into counterbored holes in the 
flange and tapped into the rear face of the tube. Re¬ 
moval was easily accomplished by inserting the 
screws into tapped holes in the flange and using them 
to back the liner out of the tube. 

A 75-mm liner assembly, made according to this 
design, was proof-fired and immediately subjected to 
a series of 50 service rounds. A second series of 300 


service rounds was then fired at a rate of 1 round per 
min. After each firing the lin^r was easily removed by 
one man in 10 min. Some of the screws had started to 
move slightly and their threads were flattened. New 
screws of harder material, equipped with lock wash¬ 
ers, were procured, and a third series of 300 service 
rounds was fired without incident. These tests dem¬ 
onstrated that the basic design of replaceable steel 
liner was sound. 

26 3 3 Successful 90-mm Liner 

Preparation of Liner Assemblies 

With the experience gained in design development 
just described, a series of three interchangeable, re¬ 
placeable, steel liners was designed and constructed 
for the 90-mm gun, Ml. This gun has extensive tacti¬ 
cal use, and, since its tube life under certain service 
conditions was quite short, the use of replaceable 
liners would materially ease the problem of supply 
and transportation. The possibility of extending the 



F igure 5. 90-mm gun tube, T19, with replaceable steel 

liner, before firing. Top: Breech-end view of tube and 
liner. Bottom: Forward end of liner, showing gas seal¬ 
ing lip. (By courtesy of Aberdeen Proving Ground.) 


CONFIDENTIAL 






508 


SHORT LINERS AND OTHER DESIGN FEATURES OF GUN TUBES 


life in this way was of especial interest in the spring 
of 1944, when consideration was being given by the 
Army to firing this gun with double-base powder in 
order to increase its muzzle velocity. 220 

The liner shown in Figure 5 was 64 in. long, and 
weighed about 250 lb, in contrast to the weight of the 
90-mm gun tube, Ml, which is 1,465 lb. This same 
figure shows the breech end of the recessed tube 
and the gas sealing lip on the forward end of the 
liner. 

For the purpose of saving time, a liner was ma¬ 
chined from a tube which had muzzle defects but a 
perfect breech end, the chamber of which was re¬ 
tained in the liner. The main tube into which the 
liner was inserted had defective rifling at the origin 
but otherwise was perfect. In recessing for the liner, 
this defect was, of course, removed. In the fabrica¬ 
tion of the three liners, the tube was used as a pattern 
in order to avoid preparing gauges. 

A liner can be assembled and disassembled by two 
men in about 20 min, which is much less than the 
time required for unscrewing the tube from the 
breech ring or for dismounting the tube and the 
breech ring. No difficulties were experienced in re¬ 
moving a liner after firing as many as 553 rounds. A 
ring of “copper” about 0.01 in. thick formed at the 
joint ahead of the liner, but always adhered to the 
liner when it was withdrawn. 


Results of Firing Tests 

The assembly with a 90-mm Ml tube and inter¬ 
changeable, replaceable steel liners was designated by 
the Ordnance Department as 90-mm Gun, T19, and 
was tested for Division 1 at Aberdeen Proving 
Ground. All tests so far have been successful. Three 
liners have been fired from the same tube. Nearly all 
rounds were fired with FNH-M2 (double-base) pow¬ 
der and APC ammunition. Firing was at normal 
rates in groups of 25 or 32 rounds. 

The first liner was fired 321 rounds and then re¬ 
placed by a second. The liner had not worn excessive¬ 
ly but was replaced on account of an increase in 
diameter of about 0.010 in. in the tube ahead of the 
liner. Star-gauge readings during the course of the 
test had indicated that wear of the forward end of the 
liner would approximately equal that of the barrel 
just ahead of it when the latter had become 0.010 in., 
so this figure was chosen as a criterion to indicate the 
time to change liners. Actually performance was sat¬ 
isfactory at this stage of erosion. 

Introduction of the second liner restored the gun to 
approximately new tube performance. This restora¬ 
tion of performance in terms of velocity is shown very 
well by the results presented in Figure 6. 

The second liner was fired 553 rounds, at which 
stage malfunctioning of the ammunition occurred 



Figure 6. Velocity loss with replaceable steel liners “A,” “B,” and “C” fired in 90-mm gun, T19. (Aberdeen Proving 
Ground Memo Reports, O.P. 6102.) 


CONFIDENTIAL 






INSERTION OF A REPLACEABLE STEEL LINER 


509 



and the test was discontinued. Its rate of erosion was 
the same as that of the first liner, as shown in Figure 
7. The third liner has been fired several hundred 
rounds. No difficulties with interchangeability or re- 
placeability have been encountered. 

In order to evaluate the potential usefulness of a 
replaceable steel liner, two further steps are needed. 
The first is to continue the firing tests already made, 
using as many additional liners as would be required 
to render the main tube unfit for further service. The 
end of its life presumably would occur because of the 
erosion ahead of the liner. The rate of this erosion is 
independent of the condition of the tube at the origin 
of rifling, for a graph of this erosion as a function of 
the total number of rounds fired in the tube, shown in 
Figure 8, does not reflect the replacement of liner 
“A” after round 321. A similar comparison after the 
insertion of liner “C” was not possible because the 
bore diameter of the tube proper decreased instead of 
increased, presumably because of coppering. No in¬ 
formation is available by which to form an estimate 


of how much erosion in this region can be tolerated 
without either seriously weakening the tube or caus¬ 
ing malfunctioning of the ammunition. 

After the true end of life had been determined in 
terms of the number N of liners that could be inserted 
in the same tube, the next step would be to find out 
whether the use of these liners was economical. The 
cost of manufacturing, and shipping to a theater of 
war, one recessed tube plus N liners should be com¬ 
pared with the total cost of manufacturing and ship¬ 
ping N monobloc tubes. A fairly rigorous comparison 
could be made on the basis of an engineering study of 
the manufacturing operation. The tolerances that 
must be maintained are very close, being of the order 
of one thousandth of an inch on the diameters of the 
liner and of the recessed hole in the tube. If there is 
more play than this small amount, the liner expands 
excessively, and then is not readily removable, as 
was found with the second design of 75-mm liner. 
With modern manufacturing methods the necessary 
tolerances can be achieved in production. 


CONFIDENTIAL 









510 


SHORT LINERS AND OTHER DESIGN FEATURES OF GUN TUBES 


- 1 -i r~ 

1 l 1 l 1 □ 

- 

□ 


□ 


□ 

- 

□ 


□ □ 

_ . O—LINER "A" 

□ 

n n □-LINER ”B" 


□ 


□ 


„ □ 


□ 

a 

) 

o 


o 

□ 

o 

o 

o 

_l_ 

- 

o 


o o o 

_n-1-1-1— 

_1_1_1_1_1_ 


0.015 


£ 0.010 


300 400 500 600 

TUBE ROUND NUMBER 


700 


900 


Figure 8. Land erosion of 90-mm gun tube, T19 just beyond junction of tube and replaceable steel liner, when fired 
with liners “A” and “B.” (This graph is based on data given in Aberdeen Proving Ground Memo Reports, O.P. 6102.) 


26 4 APPLICATION of theory of tube 

STRESSES 6 

26 4 1 Introduction 

While man}" approximate and exact theories of 
tube stresses have been formulated, the resulting 
equations (some of which are presented in Chapter 7) 
are in general too unwieldy for use in routine design 
work. For this reason, except where great accuracy is 
required, many assumptions are made, which result 
in simple equations having an accuracy usually as 
good as the fundamental data. It is the purpose of 
this section to discuss and illustrate the application 
of these simplified equations to the design of hy¬ 
pervelocity gun tubes, especially those containing 
liners. 

There are four basic assumptions made: (1) that 
the section being considered is a small section of a 
thick-wall tube having infinite length and constant 
contour, (2) that the internal pressure is equally dis¬ 
tributed and acts only radially so that there are no 
longitudinal stresses, (3) that the Lame formula ex¬ 
presses the radial distribution of tangential stresses, 
and (4) that the material is at all times in the elastic 
state. 


e A review 563 of some phases of this subject, based on work 
carried out by Massachusetts Institute of Technology for the 
Navy Bureau of Ordnance during the war, has been published 
recently. (Editor’s note in proof.) 


26 4 2 Theories of Failure 

A simple tensile or compression member can usual¬ 
ly be considered to have failed when the simple 
stress in that member reaches or exceeds the yield 
strength (or elastic limit) of the material as measured 
in the usual tension or compression test. However, 
for members subjected to more than one principal 
stress, this theory does not yield satisfactory results. 
For example, in a gun tube, failure might occur be¬ 
fore any of the principal stresses reached the yield 
strength. 

Two theories of failure of gun tubes have been ex¬ 
tensively used. Until recently, the gun designers of 
the Army and Navy used the “maximum strain the¬ 
ory” (St. Venant). This theory states that failure 
occurs when any part is strained beyond the strain 
corresponding to the yield strength as determined by 
the usual tensile test. This criterion of failure has 
been superseded by the “shear-energy theory” 
(Mises-Hencky) (also called “distortion-energy”), 
which states that failure may be considered to occur 
when the shear-energy in a unit volume of the ma¬ 
terial exceeds the shear-energy per unit volume at the 
yield point in the usual tensile test. This might also 
be expressed by stating that failure occurs when the 
“equivalent” stress exceeds the yield strength. The 
equivalent stress is calculated by equation (1), which 
is the same as equation (4) of Chapter 7. In this 


CONFIDENTIAL 






APPLICATION OF THEORY OF TUBE STRESSES 


511 


equation a e = equivalent stress; c t — tangential 
stress; <r r = radial stress; and a z = longitudinal 


stress. 

c _ — (Jr) 2 + {<Jr — (Jz) 2 + {(Jz — (JtY ( 1 ) 


Stress Distribution in a 
Thick-Wall Tube 


When a thick-wall tube is subjected to both inter¬ 
nal and external pressures, the stresses throughout 
the wall are not constant but vary from the inner to 
the outer surface, as a function of the radius. This 
distribution of the stresses is given by equations (2a, 
b, and c), which are the so-called Lame formula pre¬ 
sented in a different notation inequations (1), (2), and 
(3) of Chapter 7. Here p = Poisson’s ratio (usually 
taken as 0.3 for gun steel); p a = internal pressure; 
Pb = external pressure; r a = internal radius; r*> 
= external radius; r = radius to any section. 



For a gun tube, which is subjected to internal pres¬ 
sure only, equations (2) reduce to equations (3), 


_ Ta 2 + n 2 _ „ W 2 + 1 

<Tt V a r b 2 _ r 2 V a ff/2 _ y ( 3 a) 

(Jr = Pa. (3b) 

2 pr a 2 _2 n /0 ^ 

<Jz ^ a rb 2 — r a 2 ^ a W 2 — 1 ’ ( 3 °) 

in which IF, the wall ratio, is defined as r&/r 0 or 

db/d a . Tensile stresses are considered positive, while 
compressive stresses and pressures are considered 
negative. 

As an example, consider a tube having an internal 
pressure of 60,000 psi; inside diameter, 3 in.; and out¬ 
side diameter, 6 in. From equation (3) the three 
principal stresses are calculated as: 

(j t = 80,000 psi 
(j r = —60,000 psi 
<j z = —12,000 psi. 

Substituting these values in equation (1) jdelds a 
value of 123,200 psi for the equivalent stress. 


In order to prevent failure therefore, the material 
of which the tube is made mpst have a yield strength 
of at least 123,200 psi. This calculation does not take 
account of any safety factors, conversion from copper 
to true pressure, or proof pressure, these having pre¬ 
sumably been considered in arriving at the maximum 
pressure of 60,000 psi. 

2SAA Design Procedure for Cannon Tubes 

The Ordnance Department of the Army has devel¬ 
oped a procedure to 'be followed in the design of can¬ 
non tubes, other than auto-frettaged tubes. 293 In the 
use of this method the stress equations (3) have been 
simplified to equation (4), 

Y =wYr l v * wr + T - ( 4 ) 

for Y, the stress in pounds per square inch at which 
yielding occurs, in terms of P, the actual internal 
(powder) pressure (in pounds per square inch) and 
IF, the wall ratio of the tube, defined as above. For a 
given yield strength, the tangential resistance (max¬ 
imum pressure without stressing beyond the yield 
strength) will be R , which is defined by equation (5). 

F(TF 2 - 1) 

R = V3JF 4 + I _ ' 

R' = P' X 1.2 X 1.15 X F.S. (6) 

This tangential resistance must be not less than R\ 
defined by equation (6) in which P r = rated maxi¬ 
mum powder pressure (copper); 1.2 = conversion 
from copper gauge to actual pressure; 1.15 = ratio 
of proof pressure to rated maximum; and F.S. = fac¬ 
tor of safety. 

The factor of safety should be as high as is consist¬ 
ent with dimensional restrictions of the gun, and in 
no case less than 1.05 over the region of expected 
maximum pressure. Factors of safety at the muzzle 
(not considering muzzle bells) will normally be 
about 3. 

The yield strength measured at 0.1 per cent offset, 
as specified by U. S. Army Specification 57-106A, is 
approximately 10,000 psi higher than the true elastic 
strength of the steel. 290 Therefore the yield strength 
to be used in the above calculations for tangential 
resistance must be 10,000 psi less than that measured 
at 0.1 per cent offset. 

A report 327 has been written recently as an aid in 
the designing of guns. It contains an excellent set of 
equations, tables, and graphs covering a wide range 


CONFIDENTIAL 


















512 


SHORT LINERS AND OTHER DESIGN FEATURES OF GUN TUBES 


of wall ratios, with good explanations of the methods 
of finding the combined effects of band and gas pres¬ 
sures. There is also a well-outlined method of sep¬ 
arating the localized uniform band pressure from that 
of the traveling wave. The fundamentals of gun de¬ 
sign have been discussed in two other reports, in 
one 307 of which particular attention is paid to radial 
vibrations and in the other 310 to longitudinal stresses. 

26,4 5 Stresses in Shrunk-In Liners 

For many years both the Army and the Navy have 
made medium- and large-caliber guns of the ‘‘built- 
up” type, in which a full-length liner is shrunk into 
the tube. In general, the use of shrunk tubes and 
liners tends to reduce maximum stresses, making the 
assembly capable of withstanding higher powder 
pressures without failure than can the ordinary 
monobloc tube. Therefore in some models an addi¬ 
tional jacket is shrunk on the outside of the tube over 
the rear portion that is exposed to the maximum 
powder pressure. 

This same method is useful for preventing move¬ 
ment of an erosion-resistant liner, as has been men¬ 
tioned in Section 26.2.2. 

Since actual interference does not exist after the 
parts have been shrunk together, there is a high pres¬ 
sure between the tube and liner. This pressure creates 
stresses in the tube and liner, even with the gun at 
rest. To these stresses should be added algebraically 
the stresses produced in a monobloc tube by the 
powder pressure, in order to obtain the pressure at 
any point during firing. 

In the discussion that follows, the assumptions are 
made that (1) the radial external pressure on the liner 
is equal to the radial internal pressure on the tube, 
(2) no shear forces exist between the tube and liner, 
and (3) the external radial deformation of the liner 
plus the internal radial deformation of the tube is 
equal to the radial interference. 

For a liner with uniform external pressure, accord¬ 
ing to equations (2), the simple stresses at the out¬ 
side surface are given by equations (7a and b), and 
the unit tangential strain by equation (7c). 


\~r b 2 + r/~\ 

at L r b 2 — r a 2 J 

p b . 

(7a) 

a r = p b . 


(7b) 


1 + 

•T. * 

l 

(7c) 


in which E is the modulus of elasticity. Then the 
change in outside diameter of the liner is given by 
equation (8). 


A D l = 2 r b e t = 


2P b r b 

E l 


r b 2 + r Q 2 
7& 2 - r a 2 



(8) 


Similarly, for the tube (inside radius = r h , outside 
radius = r c and internal pressure = p b ) the change 
in outside diameter is given by equation (9). 


ADt = — 


2 p b r b 
Et 


{ r 2 + r b 2 
\r c 2 — r b 2 



(9) 


The diametral interference I between the liner 
and tube will be determined in general by equation 
( 10 ). 


I = A Dt — A Dl 


= 2 p b r b 


1 / r 2 + r b 2 

E t \r 2 ~ n 2 



+ 


1 

E l 


r b 2 + r 2 
j b 2 - r a 2 



( 10 ) 


If both the tube and liner have the same modulus 
of elasticity, E, and Poisson’s ratio, p, equation (10) 
reduces to equation (11). 


I 


2 r b I” r c 2 + r b 2 r b 2 + r a 2 
E L r 2 - r b 2 n 2 - r 2 



Pb. 


(ID 


From this equation the shrinkage pressure caused by 
any given shrink-fit interference may be computed. 
Thus, if a liner having 0.50 in. ID, 0.75 in. OD, 
E = 50,000,000 psi, p = 0.3, is shrunk in a tube 
having 0.75 in. ID, 1.3 in. OD, E = 30,000,000 psi, 
p = 0.3, an interference of 0.001 in. between the tube 
and liner will cause a shrinkage pressure of 14,800 psi. 

At the bore surface of the liner, the tangential 
stress produced by the shrinkage pressure is given by 
equation (12), which is derived from equation (2a) by 
substitution of the boundary conditions r — r a and 
p b = 0. This stress will be negative, indicating that 
the surface is in compression. When powder pressure 
is applied 


<*e 


r —r a 


— 271Tb 2 

r b 2 — r a 2 


( 12 ) 


this surface is in tension. Thus the shrinkage pressure 
tends to reduce the stress produced by the gas pres¬ 
sure, and higher gas pressures may be used than if the 
barrel had been of monobloc construction. 


CONFIDENTIAL 













PROBLEMS IN INSERTION OF EROSION-RESISTANT LINERS 


513 


26 5 SPECIAL PROBLEMS IN INSERTION 
OF EROSION-RESISTANT LINERS 

26,51 Support of Brittle Liner Materials 

The insertion of stellite liners described in Section 

26.2 was simplified by the fact that Stellite No. 21 has 
yield strength and ductility closely approaching 
those of gun steel, as may be seen by examination of 
Table 4 in Chapter 19. This is true both at room 
temperature and at the elevated temperatures that 
occur in gun barrels during firing. Also the modulus 
of elasticity and the coefficient of thermal expansion 
of Stellite No. 21 do not differ greatly from those of 
gun steel. Because of this similarity in properties the 
stress distribution in a stellite liner is not greatly dif¬ 
ferent from that in a similar steel liner. 

For a liner of a brittle material, however, a much 
less favorable situation exists. A single application of 
the powder pressure will crack such a liner inserted in 
the ways described in Section 26.2. The situation ex¬ 
pressed by equation (12) suggests a way of circum¬ 
venting this difficulty. By putting the liner under 
strong compression initially, it is possible to counter¬ 
balance the powder pressure so that even during 
firing the liner remains in slight compression. In this 
way it was possible to prevent longitudinal cracking 
of caliber .50 liners of the brittle chromium-base alloy 
described in Chapter 17. A compressive hoop stress 
of from 90,000 to 100,000 psi was imposed on the liner 
by a severe shrink-fit. 

26.5.2 Effect of Differences in 

Thermal Properties 

The successful use of a large compressive hoop 
stress in preventing cracking of a brittle liner requires 
that the coefficient of thermal expansion of the liner 
material should be not too much smaller than that of 
the steel in which it is inserted. Otherwise during 
firing, especially in a machine gun barrel, the barrel 
is likely to expand away from the liner. 

One of the difficulties of the utilization of molyb¬ 
denum as a gun liner material is that it has a low 
coefficient of thermal expansion. As maybe seen from 
the two curves in Figure 9, steel expands linearly 
more than twice as much as molybdenum. The prob¬ 
lem is further complicated by its high modulus of 
elasticity which causes it to bear an undue proportion 
of the internal pressure in the gun tube. With such a 
tube the heat exchange between the powder gases, 
the molybdenum liner, and the steel container de¬ 


pends in large measure on the ratio of the thermal 
constants of molybdenum and steel. Roughly speak¬ 
ing, the specific heat of molybdenum is one-half that 
of steel, the thermal conductivity is twice as great, 
and the thermal expansion one-third as large. There¬ 
fore, if a given amount of heat is applied to a given 
insulated mass of molybdenum its temperature in¬ 
crease is nearly twice as great as that of a similar 
mass of steel under the same conditions. For a 
molybdenum liner contained in a steel jacket and 



Figure 9. Thermal expansion of steel and of molyb¬ 
denum containing 0.1 per cent cobalt. (This figure has 
appeared as Figure 3 in NDRC Report No. A-423.) 


subjected to the hot powder gases in a gun bore, the 
temperature does not at first rise as high as if it were 
insulated, since the heat it receives is rapidly con¬ 
ducted to the steel jacket. The consequent heating of 
the steel jacket may cause it to expand so much more 
than the molybdenum liner that the latter is left 
unsupported. 

In the meantime continuing flow of heat into the 
molybdenum from the powder gases will raise its 
temperature and increase its expansion. The inter- 


CONFIDENTIAL 










514 


SHORT LINERS AND OTHER DESIGN FEATURES OF GUN TURES 


ruption of heat flow into the steel jacket would per¬ 
mit it to cool and contract until the steel and molyb¬ 
denum surfaces come into contact again, after which 
the process would be repeated. This cycle of heating 
and cooling might cause thermal stresses in molyb¬ 
denum which would cause disastrous cracking. 

In the first tests of molybdenum liners, this diffi¬ 
culty was prevented by shrinking and brazing the 
molybdenum tube into a sleeve of an iron-nickel- 
cobalt-titanium alloy, designated as Ti-Kovar. This 
alloy has the lowest coefficient of linear expansion in 
the range 20 to 500 degrees centigrade of any alloy 
that has both high strength and good elongation. 
The use of the Ti-Kovar sleeve made it possible to 
fire a 100-round burst through one of these molyb¬ 
denum liners inserted in a caliber .50 heavy machine 
gun barrel without cracking. 49 The ultimate tangen¬ 
tial strength of the molybdenum had been meas¬ 
ured as 84,000 psi. 

Its ductility was probably less than that of the 
molybdenum prepared later by the improved work¬ 
ing schedule described in Section 18.3.2. 

An answer to the question of whether the foregoing- 
situation will be of real importance in the use of a 
seamless molybdenum finer cannot be given until a 
sample of molybdenum having superior physical 
properties has been prepared. It may well be that 
molybdenum having an ultimate tangential strength 
of say 120,000 psi and slight ductility will be able to 
resist successfully the deformation just described. 
Molybdenum with this high a tensile strength would 
also have a correspondingly high compressive 
strength and would therefore stand a rather severe 



Figure 10. Cross section of a longitudinal segment or 
“stave” of a four-stave gun liner, showing the forces 
acting on it. (This figure has appeared as Figure 16 in 
NDRC Report No. A-273.) 


shrink-fit. Hence, it does not seem unreasonable to 
expect that given a seamless molybdenum tube hav¬ 
ing the properties just mentioned, it can be success¬ 
fully inserted as a gun finer by shrink-fitting alone. 

26 5 3 Stresses in Stave-Type Liners 

Another means that has been successful in virtual¬ 
ly eliminating longitudinal cracking in a molybdenum 
liner is the stave type of construction described in 
Section 18.5.2. With this arrangement the tangential 
stresses tending to produce tensile failure merely 
open up the seams between staves. The stress pro¬ 
duced in such stave-type liners has been analyzed in 
the following terms. 52 

A cross section of a stave or longitudinal segment 
for a four-stave liner is shown in Figure 10. Assuming 
that the liner fits loosely, it is acted upon by normal 
forces over the entire outside surface equal to the ex¬ 
plosion pressure p and by frictional forces pp set up 
on the outside surface by sliding of the stave in the 
barrel as the barrel expands. Neglecting bending mo¬ 
ments, the tensile stress at B, the center line of the 
outside of the stave, is given by equation (13), 

where p is the explosion pressure, p the coefficient of 
friction, a the segment angle (radians), r the radius to 
the outside of the segment, h the thickness of the 
segment. To determine the equivalent stress, zero 
axial extension was assumed so that = 0.3 (<x, 
— p). Since a r = — p, the equivalent stress, based on 
equation (1), is given by equation (14). 

<Je = 0.89 pyj^ t + 1.8~- + 1, (14) 

which may be transformed to equation (15) by com¬ 
bination with equation (13). 

< T e = 0A4p\/(par/h) 2 — 0A(par/h) + 0.8 (15) 

This equation shows that both the simple and 
equivalent stresses (for a given gas pressure) are 
a function of the nondimensional ratio ( par/In ) 
alone, and that as this ratio increases the stresses 
likewise increase. Obviously, then, it is desirable to 
make this ratio small by (1) lubricating the surface 
between the stave and barrel, (2) using a large num¬ 
ber of staves, and (3) using thick-walled staves. 

Table 1 shows the stresses in a stave-type liner for 


CONFIDENTIAL 







DESIGN OF RIFLING 


515 


Table 1 . Calculated stresses in stave-type liners for caliber .50 gun at explosion pressure of 80,000 psi. 52 


Thickness 

Coefficient 

Four-stave (a 

= t/2) 

\ - 

Two-stave (a 

= 7T) 

of liner 

of friction 

<Tt 

<T e * 

<Tt 


h (in.) 


(psi) 

(psi) 

(psi) 

(psi) 

He 

0.2 

- 51,000 

36,400 

- 21,000 

54,700 


.4 

- 22,000 

54,000 

37,000 

101,500 


.6 

8,000 

77,500 

96,000 

152,000 


.8 

37,000 

101,500 

106,000 f 

162,000f 

%2 

0.2 

- 34,000 

46,000 

12,000 

80,700 


.4 

12,000 

80,700 

105,0001 

160,000f 


.6 

58,000 

119,500 

106,000 f 

162,000f 


.8 

105,000 

160,000 f 

106,000 f 

162,000 f 


* Equivalent stress based on shear-energy theory, Equation 15. Negative signs indicate compression. Axial stress taken as 0.3 (ct< — p) corresponding 
to zero axial extension. 

f These values are practically the same as for a seamless liner. 


a caliber .50 gun at explosion pressures of 80,000 psi, 
computed by equation (15) for two different liner 
thicknesses, two different values of a , and four dif¬ 
ferent values of ju . 

An important conclusion from equation (13) is 
that, under the assumptions made, the tangential 
stress in liners of a given number of staves in guns of 
different calibers will be the same provided that the 
wall ratio and the coefficient of friction remain 
constant. 

Increasing the number of staves to decrease the 
tangential stress in each one introduces the problem 
of restraining the staves. Brazing has been success¬ 
fully used 95 for ten-stave liners fired a relatively few 
rounds in the caliber .50 erosion-testing gun (Sec¬ 
tion 11.2.1). For larger caliber guns brazing would be 
considerably more difficult and might be less suc¬ 
cessful. 

A mechanical scheme for kejdng the staves of a 
multistave liner in place has been suggested. Each 
stave is machined with a tongue on its back surface 
to fit into a corresponding groove in a steel jacket. 
The tongues should be shaped as shown in the plane 



Figure 11 . Proposed method of keying the staves of a 
liner into a steel gun barrel. (This figure was taken 
from a monthly progress report from the Geophysical 
Laboratory on Contract OEMsr-51.) 


section perpendicular to the axis represented by Fig¬ 
ure 11 so that their tightness may not be affected by 
the different thermal expansion of the stave and the 
steel jacket. In general, the back of a stave is bound¬ 
ed by four planes having a common line of intersec¬ 
tion, and similar plane surfaces form the grooves and 
thrust bearing surfaces in the jacket. Because of the 
mechanical difficulties inherent in this design, it has 
not yet been tried. 


26 6 DESIGN OF RIFLING 

26,6,1 Rifling Requirements of 

Hypervelocity Guns 

The hypervelocity guns to which Division 1 has 
given consideration have been ones for firing spin- 
stabilized projectiles. f Hence the bore surface of the 
gun tube is rifled with helical grooves in order to im¬ 
part spin to the projectile by engagement with a soft 
metal jacket on bullets or a soft metal rotating band 
on artillery projectiles. 

The rotational velocity of a projectile is directly 
proportional to the muzzle velocity when it is fired 
from guns having the same twist of rifling. Fortu¬ 
nately the stability relations, discussed in Chapter 8, 
are such that if the projectile is stable at one velocity 
it is stable at all velocities. Therefore the twist of 
rifling of a hypervelocity gun needs to be no different 
from that of a lower velocity gun in order to make 
the projectile stable. 

f The stabilization of projectiles by means of fins or in other 
ways was not investigated. It may be noted in passing, how¬ 
ever, that the erosion-resistant materials developed by the 
Division would be just as useful for hypervelocity smooth-bore 
guns. Also the absence of rifling would increase the ease of 
application and would permit the utilization of relatively thick, 
hard coatings that are not practical for rifled gun tubes. 


CONFIDENTIAL 










516 


SHORT LINERS AND OTHER DESIGN FEATURES OF GUN TUBES 


The torsional moment exerted on the engraved 
projectile band by the rifling increases as the square 
of the velocity. The strength of the band material 
cannot be increased indefinitely without seriously in¬ 
creasing the band pressure (Chapter 7). Therefore 
the strength of the band appears to be a serious lim¬ 
itation to the use of jacketed or banded projectiles at 
velocities in excess of about 4,000 fps. An answer to 
this problem is to be found in the use of pre-engraved 
projectiles, as discussed in Section 27.3. 

Regardless of whether banded or pre-engraved 
projectiles are fired, the form of the rifling of a hyper¬ 
velocity gun should be modified slightly in order to 
strengthen it, for the torque exerted on the rifling by 
the projectile also increases as the square of the 
velocity. 

26,6-2 Factors Affecting the 

Strength of Rifling 

For a given torque on the projectile, the bearing 
pressure between the rifling and the projectile is in¬ 
versely proportional to the depth of the grooves, the 
width of the rotating band, and the number of lands 
and grooves. It is obviously desirable to keep this 
pressure as low as possible to prevent wear. 

Depth of Grooves 

Increasing the groove depth in the caliber .50 ero¬ 
sion-testing gun (Section 11.2.1) from 0.005 in. to 
0.010 in. gave increased velocity-life at muzzle veloci¬ 
ties of 4,000 fps. Naturally, engraving stresses and 
radial bore pressures were increased when using en¬ 
graving-type bullets, but the resulting wear was not 
proportional. Obviously a compromise must be made 
to prevent introduction of one problem by eliminat¬ 
ing another. 

Firing tests indicated that pre-engraved projectiles 
(Section 27.3.3), having engraving twice as deep as 
normal, showed only 1 to 2 per cent greater drag than 
a smooth projectile and that depth of rifling had so 
little effect that deep rifling is not detrimental to ex¬ 
terior ballistics. Furthermore fewer and wider lands 
decrease drag by reducing the frontal area of the 
projectile, thus favoring this design. 

Number of Lands and Grooves 

The use of a large number of lands and grooves for 
high-velocity guns has been known to be undesirable 


because of erosion, since the narrow lands are easily 
crushed os worn by the engraving bullet and melted 
by the gas blast. To be sure, an increased number of 
lands and grooves increases bearing area and de¬ 
creases bearing stresses, but increased area can be 
better obtained by other means. 

Width of Lands and Grooves 

When deeper grooves are used, it appears desirable 
to make the lands wider. This reduces the bending 
stresses at the roots of the lands, and also reduces 
shearing stresses in the lands. It must be remembered 
that as the width of lands is increased, the width of 
splines engraved on the rotating band is decreased 
and may lead to excessive shearing stresses there. 
When using jacketed or banded projectiles wide 
lands require high engraving pressures and may 
require additional strength in the gun tube. This 
need not be considered with pre-engraved pro¬ 
jectiles. 

Wide lands have a further advantage in that they 
reduce the possibility of body engraving, by giving a 
larger area of support to the projectile. It appears 
desirable, when using pre-engraved projectiles, to 
have the rifling as much as two to four times as deep 
and the lands two or three times as wide as is used in 
conventional guns. It is also suggested that not more 
than 12 lands and grooves should be used in guns 
over about 3 inches caliber and a fewer number on 
smaller sizes. 

Wide lands improve launching conditions by giv¬ 
ing better support to the projectile while in the bore. 
It is easier to maintain accurate bore dimensions 
with fewer and wider lands and the effect of bore 
clearance is reduced. This is particularly important 
in minimizing yawing and balloting in the larger bore 
sizes, where the inertia of the moving mass is very 
great and the pounding produces deep and often ir¬ 
regular wear, such as has been described in Section 
10.4.10. 

Angle of Twist 

A high angle of twist is desirable for exterior ballis¬ 
tics but introduces higher side forces on the rifling, 
particularly at high gas pressures. For this reason, 
some guns have been designed with an increasing 
angle of twist toward the muzzle. This arrangement 
tends to relieve bearing stresses at the breech end but 
increases them at the muzzle. Decrease in length of 


CONFIDENTIAL 



DESIGN OF RIFLING 


517 


bearing, as the projectile goes through an increasing 
angle of twist, further increases bearing stresses. 

If exterior ballistics will permit, it is desirable, as 
far as wear is concerned, to reduce the angle of twist. 

A high velocity of rotation may also adversely 
affect launching conditions. If the projectile is unbal¬ 
anced or if there is appreciable clearance between the 
projectile and the bore, centrifugal forces introduce 
an initial yaw in the projectile. Since these centrifu¬ 
gal forces are proportional to the square of the angu¬ 
lar velocity, the value of a low angle of twist is 
apparent. 

26 6 3 R.D. System of Rifling and 
Banding 

A special design of rifling and a projectile to match 
it was proposed by Colonel G.O.C. Probert and his 
colleagues at the Research Department of Woolwich 
Arsenal. 16 This has been called the Probert system, 
or preferably the R.D. [Research Department] sys¬ 
tem. The purpose of this design was to simulate the 
bore profile of a somewhat eroded barrel and thus to 
minimize the effects of erosion and increase the barrel 


life. The design of projectile and barrel was such that 
the initial resistance of the flprmer to movement was 
as small as possible in a new gun and that the pro¬ 
portional decrease of resistance with increasing wear 
was also as small as possible. The use of this system 
in several British 3.7-in. guns proved successful. A 
trial 122 was made in the caliber .50 erosion-testing 
gun (Section 11.2.1). The projectile was made with 
two copper bands, the rear one of which had a larger 
diameter than the forward one. The rear band sealed 
the bore while the forward one was being engraved. The 
forcing cone was extended and its slope made as 
small as possible. After an erosion test of 150 rounds 
at an initial muzzle velocity of 3,650 fps using double¬ 
base powder, the increase in land diameter was not 
much less than that with ball M2, or artillery-type 
bullets, but the velocity drop was much less (170 fps, 
as compared with 320 fps). With pre-engraved bul¬ 
lets, however, the corresponding velocity drop was 
only 125 fps; consequently, no further work was done 
by Division 1 on the development of this system for 
use in hypervelocity guns. This system, however, has 
been used by the Bureau of Ordnance of the Navy 
Department in its new 3-in./70-cal. gun. 




CONFIDENTIAL 



Chapter 27 

DESIGN FEATURES OF PROJECTILES 

By F. R. Simpson a and H . L. Black h 


INTRODUCTION 

H ypervelocity guns require projectiles suitable 
for withstanding higher stresses during their 
passage down the gun bore and during flight. This is 
particularly true in the case of hollow, high-explosive 
shells. The second section of the present chapter 
summarizes what has been done in determining the 
stresses in shells and analyzing their effects on the 
design of the projectile. The most important stresses 
in the mantle of a shell are those due to band pres¬ 
sures caused by engraving during the rifling process 
(see Chapter 7), and those due to gas pressure around 
the shell body behind the band. The former, which 
usually are of greatest intensity over a short section, 
give higher local stress beneath the band at the inner 
radius of the shell than the latter. 

Two methods for reducing the radial stresses re¬ 
sulting from the band pressure on the shell beneath 
it have been investigated. The first method, described 
in Section 27.3, eliminates virtually all the radial 
compressive load by using a pre-engraved rotating 
band on the body of the shell of such design that it 
fits the gun bore and rifling snugly. This band gives 
the projectile suitable spin and satisfactory obtura¬ 
tion, and leaves only the stress resulting from gas 
pressure to be resisted by the shell wall. The second 
method, described in Section 27.4, is the use of a low- 
stress rotating band of high torsional shear strength, 
with low radial compressive resistance and reduced 
surface friction in the gun bore. 

The use of either method of reducing the radial 
load due to engraving allows either the use of higher 
gas pressure, with a corresponding increase in muzzle 
velocity, or a reduction in gun weight by decreasing 
the wall thickness and retaining the same gas pressure 
and velocity. In addition, the contribution of fric¬ 
tional wear to erosion is reduced, especially by the 
use of pre-engraved projectiles. The consequent large 
increase in barrel life resulting from the combination 

a Ordnance Research Engineer, The Franklin Institute, 
Philadelphia, Pa. 

b Technical Aide, Division 1, NDRC. (Present address: De¬ 
partment of Mathematics, Michigan State College, East 
Lansing, Michigan.) 


of pre-engraved projectiles and a chromium-plated 
gun bore is described in Chapter 31. 

Special types of projectiles—such as the sabot- 
projectiles and those for use in tapered-bore guns— 
which attain higher velocities through a reduction in 
mass of the projectile, have their peculiar problems 
of stress analysis. They are described in Chapters 29 
and 30. 

Although the testing of shell forgings for leakage 
through the base is not the immediate concern of the 
projectile designer, defects in the forging material 
which permit such leakage are serious, since prema¬ 
ture explosion in the gun may result. Two testing 
methods for revealing certain types of imperfections 
in shell bases are described in Section 27.5. 

27 2 STRESSES IN SHELLS 

27 2 1 Introduction 

The Applied Mathematics Panel of NDRC, at the 
request of Division 1, prepared a review 147 in early 
1944 of the work done in this country and in Eng¬ 
land 369 ’ 373 on the strength features of the design of 
high-explosive shells. 0 Much of the general analysis in 
that report is still pertinent, and therefore it is quoted 
with a few editorial changes (to adapt it to the present 
report) in the remaining paragraphs of this section 
and in Sections 27.2.2 and 27.2.3. In the latter section 
a paragraph has been added to bring the subject up 
to date. 

It would seem that very little consideration had 
been given to the design of projectiles for strength 
until a few years ago. Apparently the failure of cer¬ 
tain shells upon passage through the guns (with con¬ 
sequent premature bursts) led to some attention be¬ 
ing paid to the problem. It is not particularly difficult 
to make the thickness of the shell walls so great that 
no failure would occur, but it is desirable in most 
cases to make the shells rather thin in order that the 


c This review was prepared by Dr. J. J. Stoker, a research 
mathematician of the Applied Mathematics Panel, following 
visits to the persons in this country working on the problem 
and a study of their reports and those from British investi¬ 
gators. 


518 


CONFIDENTIAL 




STRESSES IN SHELLS 


519 


explosive charge carried may be as large as possible 
and also in order that proper fragmentation charac¬ 
teristics are obtained. In other words, a shell design 
is wanted which provides just sufficient strength to 
prevent failure of the shell in the gun and not much 
more. d 

The problem could be treated either experimentally 
or theoretically or by a combination of both methods. 
It has been suggested that one might design a series 
of shells of decreasing thickness, fire and retrieve 
them and then make measurements of permanent 
deformations in order to arrive at the optimum de¬ 
sign. There is much to be said for such a plan, but it 
would be rather costly and would require rather elab¬ 
orate statistical controls, since the number of vari¬ 
ables which enter into the problem is large. The 
problem has therefore been considered hitherto in the 
main as a problem in the theory of elasticity or 
strength of materials in which the stresses are to be 
calculated in the shell on the assumption of perfectly 
elastic behavior of the material, with a view to basing 
the design on the stresses thus determined. 

27,2 2 General Discussion of the 

Shell Design Problem e 

Types of Stress on a Typical Shell 

A study of the stresses developed in a shell during 
its passage through the gun requires, to begin with, 
a knowledge of the external forces exerted on it. In 
Figure 1 we indicate schematically a fairly typical 
design of a high-explosive shell. It is essentially a hol¬ 
low cylinder with rather thin walls which taper to a 
roughly conical nose at the forward end. The shell is 
closed at the rear by a base which is generally thicker 
and heavier than the side walls. The base may be 
rounded on the inside of the shell; in some cases it is 
an integral part of the shell, in others it is screwed 
into the shell. 

Shrunk into a groove cut from the outer shell wall 
is a rotating band of soft metal (copper or gilding 
metal, for example), the main purpose of which is to 
engage the rifling in the gun so that the shell will be 
set into rapid rotation about its axis as it passes down 
the gun barrel. The rotating band has an outside 

d Shells for armor-piercing purposes are an exception to this 
statement, at least as far as the forward part of the shell is 
concerned. 

e Quoted from p. 4 to 12, inclusive, of AMP Report No. 
75. 147 [See footnote (c).] 


diameter greater than the bore of the gun—an excess 
of metal is provided to ensure rotation of the shell 
and also to provide a seal which prevents the propel¬ 
lant gases back of the shell from escaping between the 
shell walls and the gun. Apart from the rotating band, 
the outside diameter of the shell is kept appreciably 
smaller than the bore of the gun except at a portion 
toward the nose of the shell (called the “bourrelet”) 
which is turned to fit rather snugly inside the gun. 
The shell is thus centered in the gun at the bourrelet 
and the rotating band. The interior hollow portion of 
the shell may be filled with various explosive sub¬ 
stances, liquid or solid. 

It is useful to consider in a descriptive way what 
happens to the shell when the gun is fired. At the 
instant of firing the very high gas pressure P g devel¬ 
oped in the chamber back of the shell gives the shell 
a very high acceleration forward. The rotating band 


ROTATING 

BASE BAND BOURRELET 



Figure 1 . Schematic diagram of a typical high-explo¬ 
sive shell, showing the forces acting on it in the bore. 
(This figure is based on Figure 1 in AMP Report 
No. 75.1) 

is “engraved” by the rifling of the gun, as described 
in Section 7.3.5, which sets up a very high mutual 
pressure, the so-called “band-pressure,” P&, in the ra¬ 
dial direction between the rotating band and the gun 
barrel as well as between the rotating band and the 
shell. f As the shell proceeds down the barrel of the 
gun, its angular velocity of rotation increases until a 
maximum is attained at the muzzle of the gun. The 
gas pressure decreases during projectile travel while 
the band pressure may or may not decrease after en¬ 
graving has been completed. 

From this description of what occurs during pas¬ 
sage of the shell through the gun, one sees that the 
external forces exerted on the shell might be classified 
into the following three types: 

1 . The gas pressure P a which acts on the base of 
the shell and also on that portion of the outer shell 
mantle which lies back of the rotating band. 


f The band pressure is considered in detail in Chapter 7. 


CONFIDENTIAL 












520 


DESIGN FEATURES OF PROJECTILES 


2 . The band pressure P b which is exerted over a 
rather narrow ring around the shell mantle. 

3. Inertia forces of various sorts arising from the 
high accelerations imparted to various parts of the 
shell as well as to the liquid or solid filling in the 
hollow portion of the shell. 8 These forces of dynam¬ 
ical origin can be further subdivided into those due 
to forward acceleration under the action of the gas 
pressure Pf and those arising from the angular accel¬ 
eration and angular velocity imparted to the shell by 
the rifling of the gun. 

Analysis of Shell Stresses 

Various methods are available for determining the 
stresses in shells. It will be helpful for our later dis¬ 
cussion to indicate in a general way how a few of the 
estimates for the stresses are made. This is perhaps 
the best way to become acquainted with the difficul¬ 
ties and complexities peculiar to the problem. 

Stresses in the Base. Let us consider, for example, 
how the stresses in the base of the shell are analyzed. 
The base is subjected on its rear face to the gas pres¬ 
sure P g and to a pressure Pf on the interior face which 
arises from the inertia of the filler inside the shell. 
The latter pressure is readily computed once the for¬ 
ward acceleration of the shell is known. This acceler¬ 
ation is found from the total mass of the shell and 
filling through the assumption that the unbalanced 
force on the shell is that due to the gas pressure P 0 , 
retarding forces arising from the driving band being 
ignored. The radial stresses in the base are then com¬ 
puted from the theory of bending of thin circular 
plates, the net pressure causing bending being given 
by -P 0 - Pf. 

Stresses in the Walls. In the more or less cylindrical 
side walls of the shell the stresses are calculated as 
follows. First of all there is a compressive stress on 
any plane section perpendicular to the axis of the 
shell which is found at any section by dividing the 
inertia force arising from the mass of that part of the 
shell (as well as the fuze, adapter, etc.), which lies in 
front of the section by the area of the cross section. 
In addition there are radial and circumferential stresses 
in the shell walls due to inertia forces arising from the 
rotation of the shell and from a liquid filling (if 
present). These forces act radially outward from the 

g The forces arising from these sources are referred to by 
ordnance engineers as forces due to “setback,” in graphic 
analogy with w hat happens to a passenger seated in a streetcar 
when the car starts up quickly. 


axis of the shell. The stresses arising from this source 
are calculated from the Lame formulas for stresses 
in thick cylinders of uniform thickness, which are 
given by equations (1), (2), and (3) in Chapter 7. 

Combination of Stresses. Once the stresses on all the 
faces of a rectangular volume element at any point 
have been determined, it is necessary to judge their 
significance in causing a possible plastic deformation 
or even a rupture. The different criteria of failure are 
discussed in Section 27.2.4. 

It might be noted that the stresses arising from the 
various types of load do not have their maxima at the 
same times: the gas and band pressures are in general 
largest during engraving, while the stresses due to 
rotation of the shell are largest at the muzzle of the 
gun. 

Limitations of the Theory 

The theory of bending of thin plates and the Lame 
theory of thick-walled cylinders are based on assump¬ 
tions which are not fulfilled in the problems consid¬ 
ered here. The base of the shell is generally so thick 
in comparison with its diameter that the theory of 
bending of thin plates is not strictly applicable. In 
addition, the boundary condition to be imposed at 
the edge of the base is rather uncertain. The Lame 
formulas are also not strictly applicable to the prob¬ 
lem of determining the stresses in the walls of the 
shell, since these formulas are derived under the as¬ 
sumptions that the cylinder is uniform in thickness, 
infinitely long, and subjected to loads which are the 
same over the entire length of the cylinder. 

In view of the complexity of the problem when 
treated as one in the exact theory of elasticity, ap¬ 
proximation formulas are the only practical ones to 
be used in carrying out stress analyses for shells, as 
discussed in Section 27.2.3. Investigators both in 
Britain 349 ’ 351 ’ 355 ’ 357 ’ 359 and in this country 34 have car¬ 
ried out a few analyses according to the exact theory, 
and compared the results with those obtained by 
using approximate methods. A fairly good agreement 
was obtained. The approximate methods tend to give 
lower values of the stresses. 

At the Catholic University a method was devel¬ 
oped of determining the stresses in a thick-walled 
cylindrical shell due to normal pressures on a rotating 
band of normal width. 34 The Ordnance Department 
extended the development by a method of applying 
the numerical results found for a given band width to 
the determination of stresses from another band, of 


CONFIDENTIAL 





STRESSES IN SHELLS 


521 


any arbitrary width, provided that the wall ratio (the 
ratio of the outer shell radius to the inner) remained 
unchanged. 546 

27.2.3 p rac tical Formulas for Shell Design 

U. S. Army Practice 

The Army Ordnance Department has devised 
means of estimating the stresses in a shell which, 
roughly speaking, consist in making use of the avail¬ 
able formulas (from the theory of elasticity or strength 
of materials) for the stresses in the various portions 
of the shell which fit the circumstances as well as 
possible. 

The principal improvement in these methods of 
calculating shell stresses, made as a result of investi¬ 
gations at Catholic University for Division 1, has 
been to take into account the effect of the band seat 
in the body of the shell, which acts like a notch as far 
as the stress distribution is concerned. The stresses in 
this case were calculated by the use of new approx¬ 
imation formulas. 17 It was assumed that a uniform 
radial pressure due to the band was applied over the 
notch and a uniform radial pressure due to the pow¬ 
der gas was applied to the mantle of the shell behind 
the notch, as illustrated by Figure 2. For the bound¬ 
ary conditions at the shoulder of the notch, radial 
displacements and their axial derivatives were fitted 
exactly; but it was not possible to calculate stress 
concentrations at the corner. The hoop stress and the 
axial stress on the inner and outer walls were ex¬ 
pressed in terms of the wall ratio of the shell, the 
width of the pressure band, h and a series of empirical 
coefficients, as shown at the end of the next section. 

Relaxation Methods * * * 5 

The British have devised a set of formulas for the 
approximate analysis of the stresses in shells, and 
they have compared the results obtained by these 
formulas with those obtained by making use of the 
exact theory of elasticity. As a consequence the prob¬ 
lem then becomes a complicated boundary value 
problem associated with a pair of linear partial differ¬ 
ential equations of second order. The fact that the 

h The static experiments described in Section 7.3 showed 

that the width of the pressure band is slightly less than the 

width of the rotating band. 

5 Quoted from p. 18 and 19 of AMP Report No. 75.1. 147 [See 
footnote (c).] 


mathematical problem is of this character and that 
the body to which it is to be applied is so complicated 
in shape explains why it is impossible to expect to 
obtain a set of formulas for the stresses applicable to 
shells of any dimensions which would be at once 
simple and exact. 

The phrase relaxation methods used so frequently 
in the British reports refers to an iteration scheme for 
solving the linear equations which arise when the 
solution of the boundary value problem is obtained 
approximately by means of the method of finite dif¬ 
ferences. These computations are very laborious and 
require a considerable amount of skill and experience 
to carry out. In addition, the solutions must be car¬ 
ried through numerically for each different shell de- 




Figure 2. The band seat in a shell acts like a notch, 
and separates the shell into three regions of stress. 


sign. The intention was to analyze a sufficient num¬ 
ber of different shells covering a fairly wide range of 
designs in order to obtain information which can be 
used in the practical design of shells. 336 ’ 338 ' 343 - 344 - 347 ’ 361 
The relaxation methods make it possible to analyze 
the effect of the band pressure on the stresses in the 
shell. In the cases analyzed it was found that the 
band pressure has, for example, a quite considerable 
effect on the stresses in the base at the regions where 
the stresses are largest. 362 The band pressure tends to 
bend the base in such a way as to relieve stresses 
caused by other loads; the reduction may amount to 
almost 50 per cent in some cases. This is probably due 
largely to the fact that the cases selected were ones 
in which the rotating band was located in part di- 


CONFIDENTIAL 






































522 


DESIGN FEATURES OF PROJECTILES 


rectly above the base. If the band were located a 
distance of a band width or two in front of the base, 
it is unlikely that such large effects of the band pres¬ 
sure on the stresses in the base would be found. Again 
it should be noted that such an analysis leads only to 
the stress distribution in a particular design and not 
to formulas for the stresses valid for any design. 

In their desire to design shells with a minimum 
factor of safety the British have carried out a few 
actual firing tests with shells designed to have ab¬ 
normally low factors of safety. The trials (for 
2-pounder shells) indicated that the factor of safety 
might be reduced quite considerably below the cus¬ 
tomary value without the risk of causing premature 
bursts. 363 

27 2 4 Criterion of Failure of a Shell 117 

It is recognized that the best criterion of plastic 
flow is that of Mises and Hencky. 514 This criterion 
has the additional advantage that its application 
does not require a determination of the principal axes 


of stress, but that any direction of axes can be used. 

According to the Mises-Hencky theory, plastic flow 
occurs if the equivalent stress, S e , which is given by 
equation (1), is larger than the yield point of the 
material in the usual tension test. 

Se 2 = §[( 66—rr ) 2 + ( rr — xx) 2 
+ {xx - 00) 2 + 4rx 2 ]. 

In this equation 66 represents circumferential or 
“hoop” stress, rr radial stress, xx axial stress, and rx 
shear stress. 

Equation (1) can be simplified for cases where the 
stress is transmitted to the shell by the rotating band. 
From symmetry, the shear stress is zero under the 
middle of the band where S e is largest. Except for 
very narrow bands, the largest equivalent stress 
occurs on the inner surface of the shell, where the ra¬ 
dial and shear stresses are zero. Equation (1) is then 
reduced to equation (2), 

S e 2 = 00^— 66xx + xx^ 

= ( 00— xx) 2 + 00 xx. 



Figure 3. Empirical function fi for computing the tangential stress at the inner wall of a shell under the rotating band. 
(This figure has been prepared from Table I in NDRC Report No. A-281.) 


CONFIDENTIAL 




























































PRE-ENGRAVING OF PROJECTILES 


523 


Summary of Method for Calculating Stresses 
in a Shell 

The procedure for calculating the stresses in a shell 
can be summarized briefly as follows. 1 

1. Calculate the average band pressure Pb as de¬ 
scribed in Section 7.3.3 by the use of P/EAI. 

2. Obtain the hoop stresses 66 1 and 66 2 at the inner 
wall, corresponding to the two cases of Figure 2, by 
the use of equations (3) and (4). 


<Wi= - 

(3) 

II 

1 

■0 

1 ^ 

I — 1 

*0 

(4) 


In these equations p' is the wall ratio (ration of outer 
to inner diameter) of the shell beyond the band seat, 
P 0 is the gas pressure, Pb is the band pressure, and 
the values of the empirical coefficients fi and / 2 are 
given in Figures 3 and 4, respectively. The hoop 
stress at the outer wall is given by equation (5), 



in which 66i, the combined hoop stress at the inner 
wall, is given by the sum of equations (3) and (4). 

3. In similar manner obtain the axial stresses at the 
inner and outer walls by means of equations (6), 
(7), and (8). 

XXl = “ ( P ' 2 - lKp' + \) Pgfi ■ (6) 

— p' 2 

xx * = - (p' 2 - i)(p' + i) P ^ 3 - 

XXe = xx£[(l — p)p' + (1 + aO/p'L (8) 

The coefficients fs and / 4 are given in Figures 5 and 6, 
respectively. In equation (8) the value of Poisson’s 
ratio p is taken as 0.3 for steel. 

4. Substitute these stresses in equation (2) to ob¬ 
tain separate values of the equivalent stress at the 
inner and outer walls. Compare these stresses with 
the actual tensile strength at the yield point of the 
material used. If the stresses at the inner and outer 
walls are both below the yield point, there should be 
no failure of the shell in firing. If the stress at the 
inner wall is above but that at the outer wall below 
the yield point, the shell should ordinarily still be 
satisfactory. If both are below the yield point, the 
design should be changed to give added strength. 


j The derivation of the equations listed in this section is 
given in NDRC Report A-281. 56 


27 3 PRE-ENGRAVING OF PROJECTILES 120 

27,3,1 Design Features of 

Pre-Engraved Projectiles 

The most promising method of reducing radial load 
on the shell resulting from band pressure is by remov¬ 
ing the usual band and replacing it with pre-engraved 
teeth or splines on the outside to fit the rifling. By 
thus eliminating the usual bore friction and the re¬ 
sulting abrasion of the bore surface, the life of the 
barrel may be increased several-fold, especially in the 
case of a chromium-plated bore surface, as is described 
in Chapter 31. The many advantages of this type of 
projectile outweigh the few disadvantages, such as 
alignment in the gun for automatic fire and the preci¬ 
sion required for engraving the teeth, both of which 
problems have been overcome in actual tests. 

Many design features of pre-engraved projectiles 
were investigated in a comprehensive ballistic re¬ 
search program undertaken by The Franklin Institute 
in cooperation with the Ballistic Research Labora¬ 
tory, Aberdeen Proving Ground, where the firings 
were made. 



Figure 4. Empirical function for computing the 
tangential stress at the inner wall of a shell due to pow¬ 
der pressure alone. (This figure has been prepared from 
Table II in NDRC Report No. A-281.) 


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524 


DESIGN FEATURES OF PROJECTILES 



1.0 1.1 1.2 13 1.4 13 1.6 1.7 1.8 1.9 2.0 2.1 2.2 2.3 2.4 2.5 2.6 2.7 23 2.9 3 JO 

FUNCTION f 3 

Figure 5. Empirical function /3 for computing the longitudinal stress at the inner wall of a shell under the rotating band. 
(This figure has been prepared from Table III in NDRC Report No. A-281.) 


The program was made extensive enough to in¬ 
clude a number of controversial points in projectile 
design that had to be settled before the use of pre-en- 
graved projectiles could be justified. 

It is evident that the flight of any projectile is 
greatly affected by its launching condition as well as 
by its performance in the air. For example, a pro¬ 
jectile with excellent exterior characteristics, such as 
low drag, will show a large dispersion if its launching 
is incorrect, thereby discounting the advantages ob¬ 
tained from its design. As indicated by the analyses 
of the experimental results set forth in Sections 27.3.2 
and 27.3.3, certain definite trends in performance are 
associated with the action of the projectile within the 
gun, and other trends are associated with its action in 
free flight. 

Recommendations 

The analysis of the results of those firing tests, 
which are discussed briefly in Sections 27.3.2 and 


27.3.3, led to the following recommendations con¬ 
cerning the design of pre-engraved projectiles. 

1. The driving teeth should be reduced to a mini¬ 
mum number and their height increased to 4 or 5 
times the depth of engraving now used on standard 
banded projectiles. It is believed that the number of 
teeth on a projectile should not exceed 12 for all pro¬ 
jectiles above 3 in. in diameter and be not less than 4 
for smaller sizes. (See Section 26.6.2 for the corre¬ 
sponding recommendation with respect to the num¬ 
ber and width of the grooves in the gun tube.) 

2. The front end of the driving teeth should be kept 
sharp radially, should be pointed circumferentially at 
an angle of about 30 degrees to the projectile axis, 
and should be flush with the body and the bottom of 
the grooves. 

3. A centering cylinder should be provided between 
the back of the ogive and the front end of the driving 
teeth for aligning the projectile axially in the gun. 

4. The origin of rifling in the gun should be just be¬ 
yond a short centering cylinder in the bore. The ends 


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PRE-ENGRAVING OF PROJECTILES 


525 


of the lands should be normal to the axis and pointed 
circumferentially at the same angle as the points on 
the projectile, in order to facilitate engagement of the 
projectile with the rifling during chambering. 

5. The front end of the driving teeth should be 
placed as far forward as possible to give engagement 
in the rifling for as long as possible, as the origin of 
rifling advances with erosion. 

6 . The length of the driving teeth should be made 
sufficient to withstand the driving torque at the high¬ 
est muzzle velocity required with the minimum en¬ 
gagement expected in a worn barrel. 

7. A smooth, narrow rear bourrelet should be pro¬ 
vided just forward of the boat-tail to give a close fit in 
the gun bore and thus insure accurate guiding. 

8 . A good method of obturating should be provided, 
such as a base cup of rubber, Lucite, or other material, 
or a narrow sealing ring of some soft material, such as 
copper, that offers very little resistance to engraving 
and consequent bore friction. 


9. For greatest accuracy and consistent perform¬ 
ance, an ogive of not less thgn 8 nor more than 10 cal¬ 
ibers should be used to give a minimum drag coeffi¬ 
cient with good stability. 

10. The center of gravity of the projectile should be 
kept well forward to give greatest stability. This ar¬ 
rangement, with the band placed well forward, brings 
the driving teeth approximately over the center of 
gravity so that the yawing tendency is reduced to a 
minimum in the gun and in flight. 

27,3,2 Interior Ballistics of 

Pre-Engraved Projectiles 

Among the factors affecting the motion of a pro¬ 
jectile within a gun, the following seem to have the 
greatest effect. (Compare Figure 14 in Chapter 10.) 

1. Yawing and balloting of the projectile within the 
bore. 

2 . Radial gas thrust against the projectile. 



FUNCTION f 4 

Figure 6. Empirical function A for computing the longitudinal stress at the inner wall of a shell due to powder pressure 
alone. (This figure has been prepared from Table IV in NDRC Report No. A-281.) 


CONFIDENTIAL 




















































526 


DESIGN FEATURES OF PROJECTILES 


3. Launching conditions. unbalanced gas pressure around the projectile, un- 

Yawing and Balloting. Excessive angular motion of balanced masses under the action of the centrifugal 
the projectile within the gun bore is detrimental to force caused by the spin, and friction forces, 
its accuracy, as evidenced by the results obtained Examination of the recovered projectiles confirmed 
when firing an eroded gun with its large bore clear- this action. There were rifling marks on the bourrelet 
ance and consequently greater freedom of radial on one side and on the rear guide on the other side, 
movement. In tests with an accurate gun, for which This indicates that the projectile followed a certain 
bore and projectile dimensions were closely controlled, set of lands spirally down the bore in a yawing posi- 
measurements were made by spark photographs of tion. As shown in Figure 8, there is good agreement 
the actual average external yaw angle for various de- among the “jump ratios” (the ratio between the 
signs of projectiles. The effect of different lengths of average external yaw angle and the calculated in¬ 
bearing between projectile and gun and of different ternal yaw angle) for the different designs, 
bore clearances—which are the most important fac- Alow inherent drag coefficient, KDo reduced to zero 
tors affecting the internal yaw angle and resulting yaw, does not seem to keep a projectile from yawing, 
flight performance of a projectile—have been deter- Conditions of launching must, therefore, be the prin- 
mined from results of these tests. cipal factor influencing muzzle jump and flight yaw- 

Figure 7 shows the type of action that may take ing. Since with the equipment available for these 
place near the muzzle of a gun at the time of launch- tests, the muzzle jump could be compared only qual- 
ing. As the projectile emerges from the bore, it may itatively, only the average yaw angles of the various 
quickly develop yaw, resulting from a combination of projectiles observed can be compared. At time of 


GUN MUZZLE 



GUN MUZZLE 




EXTERNAL YAWING 


Figure 7. Relation of internal action to yawing of projectile in flight. 

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PRE-ENGRAVING OF PROJECTILES 


527 



Figure 8. Relation of interior to exterior yaw. 


launching, there is about a thirteen-fold increase in 
yaw angle for most pre-engraved projectiles, as com¬ 
pared with a twenty-fold increase for artillery type, 
and a thirty-fold increase for one with a narrow for¬ 
ward band. 

Gas Thrust. The behavior of the forward-banded 
' type of projectiles represented by design No. 8, Fig¬ 
ure 8, indicated that an unbalanced radial thrust 
from the pressure of the powder gas takes place in the 
gun bore, pushing the base or even the whole projec¬ 
tile against the rifling on one side. Because of the 
large pressure area on the base of this type projectile, 
there must be an excessive gas kick at the instant the 
bullet leaves the muzzle. A certain amount of rear 
guiding and obturation probably would reduce the 
muzzle jump and exterior yaw, and at the same time 
give excellent flight characteristics to the bullet, sim¬ 
ilar to those exhibited by design No. 1. 

Launching Conditions. Results of firing the artil¬ 
lery-type projectile (design No. 7, Figure 8) indicate 


that this type has a greater average exterior yaw than 
the pre-engraved type. An explanation of this is hard 
to make. While the form factors and drag are similar, 
the conditions at time of launching may be such that 
there is excessive disturbance at the muzzle due to the 
large gas pressure drop around the projectile causing 
high velocity gas jets that develop unbalanced forces 
when they hit any bullet irregularity. Design No. 7 
has a cannelure in the band and a slight change in 
contour near the bourrelet, both of which are not 
present in design No. 3. In design No. 1 with the two 
bands there is a sealing effect around the body of the 
projectile caused by the presence of an expansion 
chamber, as it were, between the front and rear bands. 
This cuts down the jet effect because of the lower 
pressure-drop and reduces the muzzle disturbance, 
since there are no irregularities forward of the gas 
jets. Irregular friction forces at launching of design 
No. 7 may also account for some of the additional 
yaw. 


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528 


DESIGN FEATURES OF PROJECTILES 


27 3 3 Exterior Ballistics of 

Pre-Engraved Projectiles 

The exterior ballistics of pre-engraved projectiles 
were carefully investigated to make certain that no 
feature of their design would be detrimental to their 
flight. Careful checks were made of each design vari¬ 
able by changing only one feature at a time. Extreme 
care was used during the firing tests to duplicate fir¬ 
ing conditions, spark photographs, and analysis of 
performance. The consistency of the results shows 
that they are dependable to within about 1 per cent. 

The following design variables were checked with 
groups of similar projectiles and the observed data 
were evaluated. 

Number of Rotating Bands. (Group 1, Figure 9.) 
From a comparison of the three projectile types in 
this group it has been possible to evaluate the effect 
of adding each band or circular obstruction to a high- 
velocity projectile. Comparison of the drag of the 
single-banded projectile with the smooth one shows 
an increase of 4 per cent, while the addition of another 
band, even though its sharp, front edge is at the point 
of tangency of the ogive, adds only 1.7 per cent addi¬ 
tional drag. Spark photographs of this projectile 
(Ll-107) in flight confirmed this result. The extent to 
which the drag of the rear band is greater than that 
of the front band indicates that the region between 
the base of the ogive and a point back of the center of 
gravity is not so sensitive to drag as is the base or 
boat-tail region, particularly on a projectile of this 
type with a short ogive radius. 

Position of Bands. (Group 2, Figure 9.) The con¬ 
clusions reached from a study of group 1 with regard 
to the relative effect of changes made on the base por¬ 
tion of the projectile and on the portion just back of 
the ogive, were confirmed by results from group 2 al¬ 
though they were not so pronounced. This probably 
can be accounted for by the absence of the low-pres¬ 
sure region in either type, with a general lessening of 
disturbances to the boundary layer. In this connec¬ 
tion, group 7 of Figure 9 should be noted because of 
its comparison of a boat-tailed and a flat-base pro¬ 
jectile. There the effect of changes is more pronounced, 
both because of the increased moment of the disturb¬ 
ance and because of the re-entrant stream lines being 
more effective in producing disturbances. 

Width of Bands. (Group 3, Figure 9.) In this series, 
the rear edge of all the driving bands was kept in the 
same position with reference to the boat-tail and the 
band was widened by adding to the front portion. 


Again there is indicated a tendency for the front edge 
of the driving band to have less effect on the drag as 
it is placed closer to the ogive. But what is perhaps 
more important with reference to pre-engraved pro¬ 
jectiles is a complete absence of any indication that 
the deep engraving (0.010 in. is twice the standard 
depth of 0.005 in.) on a wide band has any noticeable 
effect on the drag. It indicates that there is plenty of 
opportunity to increase the depth of the engraving to 
give very low unit pressure on the sides of the rifling 
in a gun without appreciably affecting the drag, as 
advocated in Section 26.6.2. This will allow greatly 
increased muzzle velocities with increased accuracy- 
life of the gun. 

Leading Edge of Bands. (Group 4, Figure 9.) In or¬ 
der to find the effect of pointing the leading edge of 
the splines on pre-engraved projectiles, which is one 
way of indexing them for automatic firing (Section 
28.4), a comparison test was made of the drag with 
the front edges of the bands square, conical, and 
pointed circumferentially. As was expected, the square 
edge at the ogive used on the Ll-107 type gave the 
highest resistance, but when the splines were pointed 
and the front edge set slightly back of the ogive, the 
drag became exactly the same as that of a narrow 
band whose front edge was even further back and cut 
conical. It is desirable to have the front edge of the 
driving bands on pre-engraved projectiles square in 
order to center them when entering the bore. The 
fact that pointing of the lands cuts down the drag as 
much as 3 per cent indicates that the proposed design 
of pre-engraved projectiles should have excellent 
flight characteristics. 

Radius of Ogive. (Group 5, Figure 9.) The effect of 
lengthening the radius of ogive is clearly demonstrated 
in group 5. The projectile with a 19-caliber radius of 
ogive shows a drag coefficient less than half that for 
the one with a 5-caliber radius. However, both these 
designs show a dangerously low stability factor, while 
the bullet with an intermediate radius shows good 
stability although a higher drag than the one with the 
long radius of ogive. This indicates the importance of 
balancing all the factors so that a high-velocity pro¬ 
jectile will give low drag but, at the same time, be de¬ 
pendable in flight. This bullet had a low yaw angle 
and was otherwise quite dependable, even with a flat 
base which might have been boat-tailed to reduce 
further the drag. 

While it is desirable to have a long nose radius, it is 
important that the length of guide of the body of the 
projectile shall not be shortened so much as to handi- 


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PRE-ENGRAVING OF PROJECTILES 


529 


cap the launching. This might also throw the center 
of gravity too far back from the nose of the projectile, 
resulting in poor stability. 

Location of Center of Gravity. (Group 6, Figure 9.) 
The effect of changing the location of the center of 
gravity of a projectile is shown by the three examples 
in group 6. In one the center of gravity is located well 
forward in the smooth projectile, C-43; at an inter¬ 
mediate point in a wide-banded projectile, C-30; and 
well back in the third one, C-34. The stability factor 
is high with the first type and very low for the last, 
but the form factors are reversed accordingly. 

Types C-30 and C-34, with the same drag factor 
0.132 and same contour, show the effect on the form 
factor of moving the center of gravity toward the rear 
a distance of 0.212 in. This was accomplished by 
making the nose of the C-34 projectile of aluminum. 
Their general performance was much the same but 
the second was less reliable in flight. 

Boat-Tailing. (Group 7, Figure 9.) A good indica¬ 
tion of the effect of boat-tailing a flat-based bullet is 
shown in group 7 by comparison of two otherwise 
identical projectiles: C-33 (boat-tailed) and C-31 
(flat-based). Here the change from a flat base to a 
boat-tail reduced the drag by more than 2 per cent. 
At the same time the yaw in flight was also reduced 
by this streamlining, indicating a tendency of the 
boat-tail to stabilize the flight. 

Figure 10, which illustrates the effect of various de¬ 
sign factors on the reduction of drag, relates this 
tendency to the thickness of the boundary layer with¬ 
in the streamlines surrounding the projectile and 
leading to vortices set up at the corner of the flat 
base. It indicates that, for high velocities and long 
radii of ogive, a boat-tailed base is very important as 
is any change that will make the rear portion as 
smooth as possible. Here, as in air-foil design, any 
tear-drop section such as the Joukowsky profile is an 
excellent one for obtaining low drag. 

Degree of Smoothness or Reynold's Number. (Group 
8 , Figure 9.) The three projectiles in this group, all 
having nearly the same form factor and physical di¬ 
mensions, demonstrate the reduction in drag that can 
be obtained by removing roughness of any kind that 
tends to interfere with the laminar flow of the stream¬ 
lines around the projectile, as visualized in Figure 10. 
The removal of one of the two bands from projectile 
Ll-107 gives the single-banded type C-39, with a 3 
per cent lower drag. Removal of the single band to 
give a perfectly smooth projectile, C-43, reduces the 
drag 2J^ per cent more. 


The photographs of these three projectiles in flight 
showed that the one with t\yo bands, Ll-107, had a 
heavy shock wave coming from the front and rear 
edges of each band, in addition to the head wave. The 
second, C-39, had the front band removed; the heavy 
wave from the front edge of the first band disappeared 
but a light wave still persisted from a shoulder 0.002 
in. high radially near a point at which the rear of the 
front band was located in Ll-107. The waves at the 
rear band were about the same as before but, due to a 
chamfer on the front edge, the shock wave was flat¬ 
tened out smoothly. 

Lastly, the smooth projectile C-43 showed no heavy 
shock wave except that at the head. This should, 
therefore, be the source of most of the drag. 

As far as an actual change in the surface smooth¬ 
ness or Reynold’s number is concerned, it is difficult 
to obtain much of an improvement in drag simply by 
polishing a bullet, although tool marks and surface 
irregularities must be avoided as much as possible. 

The “paddle-wheel” action of the deeper driving 
splines used on pre-engraved projectiles gives no in¬ 
dication, in the spark photographs, of causing any- 
appreciable drag, because their action probably takes 
place in a region of turbulence and recirculation 
created by the front edge of the band. Consequently, 
the effects of fringing at the rear edge of a self-en¬ 
graving band should not increase the drag greatly 
unless it is near the boat-tail. Irregular fringing, how¬ 
ever, will tend to cause rotary unbalance of the bullet 
in flight, with resulting large dispersion. 

Conclusion. The splines on pre-engraved, high-ve¬ 
locity projectiles are no more detrimental to the 
flight of a projectile than on any standard banded 
type. Furthermore, their location can be placed well 
forward of the base of the projectile to give a decided 
ballistic advantage and prolong the engagement life 
of a gun. 

It is interesting to note that reducing all the rough¬ 
nesses and obstructions on Ll-107 (group No. 8) re¬ 
duced the drag factor from 0.136 to 0.129, whereas 
changing the radius of ogive from 5 calibers to 19 
calibers (group No. 5) reduced the drag coefficient 
from 0.132 to 0.065. This fact indicates that rotating 
bands and pre-engraving of projectiles affect the drag 
very little in comparison with the shape of the ogive. 

27 3 4 Manufacturing Techniques 

The pre-engraved caliber .50 projectiles used by 
Division 1 have been made by forcing soft steel 


CONFIDENTIAL 



530 


DESIGN FEATURES OF PROJECTILES 



Figure 9. Effects of design on exterior ballistics of pre-engraved, caliber .50 bullets grouped according to design para¬ 
meters. 


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PRE-ENGRAVING OF PROJECTILES 


531 



Figure 9. ( Continued) 


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532 


DESIGN FEATURES OF PROJECTILES 


VORTICES OR EDDYS DEVELOP AT 
FRONT EDGE OF BAND AND AT 
CORNER OF SQUARE BASE 



PADDLE-WHEEL ACTION OF ENGRAVING 
SHOWS NO INDICATION OF AFFECTING 
BOUNDARY LAYER 


B = THICKNESS OF BOUNDARY LAYER. TEST SHOWS 
.2 50" AT M* 3.2 FOR 5 CAL OGIVE 

STREAMLINES 

SHOCK WAVES OR STREAK LINES 

ESSENTIALLY A STABLE DESIGN 


SHORT RADIUS GIVES MORE DRAG 
BECAUSE D s OF STREAMLINE ENVELOPE 
OR BOUNDARY LAYER IS LARGE 


FLAT-BASED TWO'BANDED SHORT RADIUS OF OGIVE 



INCLINED TO BE UNSTABLE IF 
RADIUS OF OGIVE IS TOO LONG 


LONG RADIUS GIVES SMALLER D $ 
BUT REAR BAND AND FLAT BASE 
GIVE ADDED RESISTANCE TO 
COUNTERACT ADVANTAGE OF LONG 
OGIVE 


FLAT-BASED TWO-BANDED LONG RADIUS OF OGIVE 


VORTICES AT FRONT EDGE OF 
REAR BAND CAUSE RESISTANCE 


BOAT-TAIL REDUCES RESISTANCE 
OF BULLET TO REENTRANT STREAMLINES 



BOAT-TAILED TWO-BANDED LONG RADIUS OF OGIVE 


SMOOTH OUTLINE AT REAR ELIMINATES 
VORTICES REDUCING REYNOLD'S NUMBER 
AND GIVING STREAMLINE FLOW 



B 



WITHOUT 


BOAT-TAILED AND WITHOUT REAR BAND 


Figure 10. Design factors in the reduction of drag for pre-engraved bullets at hypervelocity. 


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LOW-STRESS ROTATING BANDS 


533 


blanks through a stepped series of dies, used either 
individually in a punch press or in series in a special, 
long-stroke broaching machine. These dies make a 
series of grooves in the band which are accurately 
concentric with the bourrelet and rear guide, have the 
same twist as the rifling, and are of the correct size to 
fill all the grooves completely. To obtain this accu¬ 
racy, it is necessary to use guides that center the band 
on the projectile blank when it enters each cutting die. 

After engraving, all the teeth are pointed simul¬ 
taneously on the front end in one operation using a 
special machine that has been developed for generat¬ 
ing by a helix the proper cylindrical surface, accur¬ 
ately centered and matched to the engraving within 
0.001 in. There is no doubt that this can be done com¬ 
mercially on a large scale. By Parco-Lubrizing k these 
steel projectiles, rusting is avoided and a low friction 
means is provided for sliding through the gun. 

By using a special adapter, a portion of the point¬ 
ing machine is also used in pointing the rifling in the 
gun. This requires changing to another type of cutter 
and, by using the same machine, the same angle of 
pointing is generated so that there is plane contact 
with the points on the projectile. The ratchet grooves 
in the breech of the barrel are cut in a jig to position 
the rifling accurately in relation to the firing mechan¬ 
ism. The relative cost of these changes on the gun 
barrel are slight, but are proportionately greater on 
the projectiles. 

The manufacturing techniques for the 37-mm gun, 
T47 and projectiles (see Section 31.7) were essentially 
the same as those for the caliber .50, except that 
everything was done on a larger scale. Larger dies 
were required and consequently a heavier press had 
to be used. A bigger machine was necessary for the 
greater amount of metal removed from the teeth in 
the pointing operation, although the amount of cut 
could be adjusted. The essential differences consisted 
in more machining being required on the base of the 
cases and in necessary alterations to the firing mech¬ 
anism. The latter changes are described in Section 
28.4.4. 

27 4 LOW-STRESS ROTATING BANDS 

If pre-engraved projectiles are not to be used, it is 
very desirable to design rotating bands giving low 

k Parco-Lubrizing is a trade-marked process (Parker Rust 
Proof Company, Detroit, Michigan) for applying to steel a 
wear-resistant coating that consists chiefly of an admixture of 
iron and manganese phosphates. 


stresses. Accordingly, studies for this purpose were 
made at The Franklin Institute, the Leeds andNorth- 
rup Company, and the Catholic University of Amer¬ 
ica. i 

27 4 1 Caliber .50 Projectiles 

Firing Tests 

Banded caliber .50 bullets of the seven classes listed 
below were tested to determine the effect of rotating- 
band design on the force required for engraving. 

1. Regular artillery-type (see Figure 3 of Chapter 

11 ). 

2. Artillery-type with obturating ring of 0.530 in. 
diameter, and with band diameters varying from 
0.493 to 0.510 in. 

3. Artillery-type with radial holes, 0.055 in. in dia¬ 
meter drilled in rotating band. 

4. Artillery-type having all-steel rotating band 
with six circumferential grooves 0.0155 in. in depth. 
(Figure 11). 

5. Types 3 and 4 with holes or grooves filled with 
solder. 

6. Several of foregoing types with cadmium plate 
0.003 in. thick on the band. 

7. Pre-engraved, double-banded steel (see Figure 
3 of Chapter 11.) 

The erosion-testing gun, described in Section 11.2.1, 
was used in these tests. 122 The barrel was of mono¬ 
bloc gun steel construction with grooves 0.010 in. in 
depth. Precautions were taken to eliminate as far as 
possible all variables that would affect the maximum 
powder pressure, except the band design itself; the 
average maximum pressure attained on firing 6 rounds 
was then taken as a measure of the engraving force. 

Reduction of the band diameter (class 2) gave a 
consistent decrease in powdor pressure; the smallest 
band noted, of diameter 0.493 in., gave a pressure of 
41,590 psi as compared with 53,950 psi given by a 
0.510-in. band; this represents a pressure decrease of 
23 per cent. 

Little improvement, only about 5 per cent in the 
case of 81 holes, resulted from drilling holes in the 
band (class 3). However, when these were filled with 
solder (class 5), the reduction in pressure was of the 
order of 9 per cent for bands with either 54 or 81 holes. 

The grooved steel projectile, either plain (class 4), 
or with cadmium plating (class 6), or filled with solder 
and cadmium-plated (class 6), diminished the powder 
pressure by approximately 7 per cent. 


CONFIDENTIAL 




534 


DESIGN FEATURES OF PROJECTILES 


to FINISH ALL OVER CONCENTRIC 

o ^ AND TO WEIGHT LIMITS 

+i* 



Figure 11. Drawing of artillery-type, caliber .50 bullet having an all-steel rotating band with six circumferential grooves. 


The completely pre-engraved bullet (class 7) gave 
by far the best results, with pressures reduced by a 
third (see section 27.3). 

Static Tests 

In addition to the firing tests, static tests were con¬ 
ducted to determine the force required to engrave 
several of the projectiles listed above. It was found 
that little reduction was brought about by the use of 
plain bands with drilled holes, but when the holes 
were solder-filled the force required was diminished 
by 70 per cent. Cadmium plating diminished the 
static load by 25 per cent. 

Erosion Test of Projectiles with Grooved 
Steel Bands 

Particular interest was attached to the grooved 
steel-banded projectile (class 4) because it could be 
converted readily from standard-type bullets already 
in production. From the point of view of lowered 
radial stresses only, these projectiles should have been 
better without the solder filling. However the rather 
narrow circumferential fins would have been sub¬ 
jected to large bending moments. The lead-tin alloy 
which filled the grooves served to reinforce the fins 
and also to form a lubricant “in situ.” The entire 
bullet was then cadmium-plated to reduce friction. 
Static tests had indicated a 50 per cent reduction of 
engraving force from that required for the standard 
artillery type. 


This design, shown in Figure 11, seemed very 
promising, and a number of these projectiles were 
fired in an erosion test. It was found that land and 
groove erosion at the origin of rifling, and also muz¬ 
zle-velocity drop, were no greater than those ob¬ 
served with a copper-banded bullet. Land erosion 
from 10 in. beyond the origin of rifling was, however, 
greater. Recovered bullets, shown in Figure 12, illus¬ 
trate normal engraving and failure to engrave caused 
by wear in the gun. 

27 4 2 Larger Projectiles 114 

While the stresses can be lowered by making either 
the width or the diameter of the band smaller, the 
functions of the band set lower limits on these quanti¬ 
ties, particularly in the case of a worn gun. In order 
to impart spin to the projectile, a minimum area of 
band metal must be in contact with the driving face 
of the lands; and similarly, satisfactory obturation 
demands sufficient metal to fill the grooves. 

37-mm Projectiles 

Effect of Cannelures. A natural approach to the 
problem of reducing stresses is to consider the effects 
of grooves, or cannelures, in the band. The Service 
design for the 37-mm shot, M51, has a cannelure ap¬ 
proximately 0.1 in. wide and 0.022 in. deep. This 
depth is great enough so that the lands of the gun 
should not touch the bottom of the cannelure. After 
engraving, however, enough metal is found to have 


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LOW-STRESS ROTATING BANDS 


535 


been pushed into the cannelure to fill it to the level 
of the rest of the engraved portion. This metal in¬ 
creases the radial and axial loads without helping 
support the side thrust necessary to cause.rotation. 
The addition of a number of similar cannelures would 
in effect only shorten the band; the limits would still 
apply; and the overall result would be to add metal 
without increasing the effectiveness of the band. 

If instead the band groove is deepened, the result 
of plastic flow will be to force metal into the bottom 
of the cannelure, out of contact with the gun lands, 
thus decreasing the radial load and friction. Addition 
of other cannelures of this type will further reduce 
stresses. 

It has been shown 64 - 114 ’ 372 that the extruded fringes 
of band metal increase radial load without contribu¬ 
ting to the lateral bearing surface between the face of 
the gun lands and the projectile grooves. The projec¬ 
tile body has a groove adjoining the rear of the rotat¬ 
ing band which, however, is not large enough to pre¬ 
vent the formation of fringes. To correct this the body 
groove was enlarged. 

Experimental Tests. The theory was then subjected 
to experimental check by the method described in 
Section 7.3. The original cannelure of depth 0.022 in. 
was first replaced by one of 0.040 in. with marked 
improvement. There was evidence that even deeper 
cannelures were desirable, and the final tests were 
performed on bands with a cannelure depth of 0.065 
in. The groove in the body back of the band was of 
such a depth (approximately 0.03 in.) that the bottom 
of it was level with the bottoms of the cannelures. 
With a single such cannelure and the body groove, 
both radial and axial loads are reduced to about 70 
per cent. Addition of a second deep cannelure cut the 
loads to about half their normal value for a band of 
this size. Since the effective length of the band has 
been diminished, the second cannelure need not be 
deep to cause a further reduction in stress. The resist¬ 
ance to side thrust would in the latter case be reduced 
by approximately 20 per cent. These reductions have 
been confirmed in firing tests at Aberdeen Proving 
Ground. 

In the estimates made for radial loads in the var¬ 
ious designs of bands, the theory of Effective Area of 
Interference [EAI] developed in Chapter 7 was used. 
The method was readily applicable in the earlier tests. 
With deeper cannelures, however, it was necessary to 
introduce negative terms for the portions of the can¬ 
nelures and body grooves having smaller diameter 
than the gun lands. The calculated negative contri¬ 


butions turned out to be about twice too large; this 
indicates that the space prdvided to accomodate the 
extruded metal is only half used. 

Cadmium Platipg. Further experiments were con¬ 
ducted to study the effect of cadmium plating of the 
band. The plate applied was 0.002 in. thick. The 
radial load was not affected, but the axial load was 
reduced by a factor 2 or 3. The overall effect of the 
new design, with cadmium plating, was to produce a 
band which has a radial load about half and an axial 
load about one-fifth to one-sixth of the loads for a 
normal band of the same dimensions. 



Figure 12. Normal engraving and failure to engrave; 
shown by collected artillery-type caliber .50 bullets with 
all steel rotating bands. (A) Round #156: Bullets skid¬ 
ded along bore; band diameter reduced to 0.492 in. 

(B) Round #83: Normal band engraving. 

75-mm Projectiles 

A 75-mm AP projectile, M72, with the rotating 
band modified by a cannelure 0.098 in. wide and 
0.070 in. deep was tested for radial load. This gave 
20,600 lb/in. circumference in a tube section of wall 
ratio 1.75 with French-type rifling, in which the 
corners of the lands are rounded. If the band had 
been of equivalent length, without a cannelure, a 
radial load of about 40,000 lb/in. circumference 
would have been expected; for an unmodified Service- 
design band, the load would have been about 53,000 
lb/in. circumference. 

A comparison of band stresses during firing was 
made for a 75-mm, AA gun T22 at Aberdeen Proving 
Ground. This gun has rectangular rifling, as con¬ 
trasted with the French-type rifling used in the gun 
mentioned in the previous paragraph. HE shells were 
used. The Service-type band has an overall length of 


CONFIDENTIAL 




536 


DESIGN FEATURES OF PROJECTILES 


0.86 in., the forward 0.14 of which tapers at 8°30'. 
The modified band has a cannelure of width 0.1 in. 
and depth 0.05 in. near the center of the flat section. 
As determined by a gauge about 25 in. toward the 
muzzle from the origin of rifling, the radial load 
averaged 43,200 lb/in. circumference for the standard 
bands, and 34,700 lb/in. circumference for those with 
a cannelure. The corresponding values of P /EAI are 

40.5 and 40.4. The agreement is remarkable. 

It would appear desirable to carry out firing tests 
that have been suggested 117 to gain further informa¬ 
tion about the effects of deep cannelures of rectangu¬ 
lar cross section. In particular, it should be determined 
whether erosion is decreased and whether bands that 
have been so modified seal satisfactorily in eroded 
tubes. 

British Investigations 1 

The British have carried out experimental and 
theoretical studies on the subject of the proper design 
of the rotating band itself. 362,372,388,556 One of their 
ideas is to separate the two functions of the band, i.e., 
that of causing rotation of the projectile and that of 
acting as a seal to prevent the escape of the propellant 
gases past the shell. Since these are two very different 
functions, it is not at all unlikely that a design which 
separates the two might lead to a band which would 
cause a much smaller band pressure. One very interest¬ 
ing experiment was carried out with a double band 
on a 17-pounder shell; the front band was turned 
down so that there was no excess copper and the rear 
band was modified so that a seal would be ensured. 
The strain gauge readings on the gun when the shell 
was fired indicated that the band pressure was negli¬ 
gible. In addition, the stresses in an experimental 
band design have been calculated using relaxation 
methods. 359 

27.5 METHODS FOR DETECTING 

DEFECTS IN SHELLS 

Emphasis on accurate determination of the stresses 
in a shell mantle is important to allow the designer to 
provide sufficient metal in the walls to prevent yield¬ 
ing. But it is also necessary that the metal used be 
homogeneous and free of slag inclusions, pipes, cracks, 
or other defects that may cause failure in service by 
the hot gases seeping through the base or the walls of 

1 This paragraph is quoted from p. 20 of AMP Report No. 
75.1. 147 [See footnote (c).] 


the shell and igniting the explosive within. For this 
purpose, two methods have been developed for test¬ 
ing the finished shells to detect defective openings. 

27 5 1 First Testing Method 

Ammoniated water is used as a testing medium. 
The liquid is forced against the base of a shell forging 
under a pressure somewhat greater than maximum 
gas pressure that will be encountered in a gun. An 
indicator, such as moist litmus paper, is placed inside 
the shell cavity. Defective openings in the base of the 
shell are revealed by a change in color of the indicator 
caused by fumes from the ammoniated water. 

The apparatus used in this first method employs a 
special pressure plate for a hydraulic press. In this 
plate the base of the shell fits tightly around the rim, 
sealing a cavity beneath, which connects to a high- 
pressure tank and compressor. Inside the shell is a 
steel form against which the ram of this quick-act¬ 
ing hydraulic press applies the necessary load for 
sealing the base. The valve from the compression 
chamber is then opened for a brief time, after which 
it is closed and the ram lifted. If there is a defect in 
the shell base that would allow the slightest amount 
of gas to get through, the indicator changes color. 

This method is very fast and suitable for produc¬ 
tion line testing. However, there are certain disad¬ 
vantages, one of which is the presence of ammonia 
fumes in the air surrounding the equipment that may 
affect the operators and the surrounding equipment. 
However, this disadvantage can be offset, to a large 
extent, by the use of a greased-paper seal at the top of 
the forging, and by the use of ventilating fans in the 
room. Another objection is the fact that this equip¬ 
ment is more expensive than the second method de¬ 
scribed below, although its cost is not high when it is 
considered that one set of equipment would take care 
of several production lines that were located in the 
same or nearby buildings. 

27,5,2 Second Testing Method 

Another method is a self-contained apparatus for 
use under a hydraulic press. It consists of a stationary 
piston in a movable cylinder supported on springs, 
sufficiently powerful to hold a seal on the base of the 
shell before compression of the liquid begins. The 
apparatus is contained in a guide cylinder on which 
an indicator is mounted for measuring the relative 
travel of the moving cylinder carrying the shell. The 


CONFIDENTIAL 





METHODS FOR DETECTING DEFECTS IN SHELLS 


537 


space above the top of the piston and surrounded by 
the movable cylinder is completely filled with liquid, 
such as water or Varsol (a high-grade kerosene). 

This method is slower and less sensitive than the 
first method and, therefore, is not capable of such fast 
production. It is well suited for use in a shop where 
chemical fumes are not desirable. 


Several sets of tests were conducted on 75-mm and 
3-in. shells, with Varsol used as the liquid. The results 
indicated that a shell will be satisfactory and without 
defects through the base up to about 2.35-mm deflec¬ 
tion of the indicator. From 2.35 to 3.00-mm deflection 
the shells are doubtful and above that value they 
must be rejected. 


CONFIDENTIAL 



Chapter 28 

AUTOMATIC GUN MECHANISM 

By William H. Shallenberger a 


28i INTRODUCTION 

I n addition to making hypervelocity guns practical 
by the control of erosion, considerable work is re¬ 
quired to make them practical from a mechanical 
standpoint. That is, a mechanism designed for medi¬ 
um-velocity weapons may not be entirely suited to 
hypervelocity guns. 

One project undertaken by Division 1, NDRC, was 
the development of a 20-mm automatic aircraft can¬ 
non. Although this gun was designed to use standard 
ammunition rather than special hypervelocity ammu¬ 
nition, it had certain characteristics that made it 
particularly suited for use with pre-engraved projec¬ 
tiles (Chapter 31) and hypervelocity ammunition in 
general. Therefore, discussion of that particular gun 
makes up a major portion of this chapter (Section 
28.2). It describes in a general way the difficulties 
encountered in operation and means whereby the dif¬ 
ficulties were overcome, and presents certain conclu¬ 
sions and recommendations for the use of future de¬ 
signers of such weapons. Section 28.3 treats of some of 
the problems involved in designing an automatic gun 
mechanism for use with hypervelocity ammunition. 
Loading and indexing mechanisms for pre-engraved 
projectiles are discussed in Section 28.4. 

The information for this chapter is taken largely 
from published NDRC reports, which cover the work 
in more detail than can be given here. Some of the 
ideas presented herein, however, have not been pub¬ 
lished elsewhere, to the author’s knowledge. 

28 2 JOHNSON 20-mm GUN b 

28 2 1 Introduction 

In the summer of 1942, the Navy desired to in¬ 
crease the firepower of aircraft by replacing the cali¬ 
ber .50 Browning machine gun with the 20-mm AN- 
M2 (Hispano-Suiza) automatic cannon. The Navy 
Department requested Division 1, NDRC, to develop 

a Engineer, Engineering and Transition Office, NDRC. 
(Present address: Granada Hill, San Fernando, California.) 

b This section has been condensed from an NDRC report 126 
by the same author. 


an integral belt-feed for this gun, but a preliminary 
examination of the problem indicated that it would 
be preferable to develop an entirely new gun than to 
try to make the necessary modifications in an existing 
gun. To this end, Division 1, NDRC, undertook Pro¬ 
ject NO-124 to design and build a firing model of a 
20-mm automatic aircraft cannon having an integral 
belt-feed. Under this project 0 three models were 
designed. 

The first model, known as Model I, was designed 
and built to test the basic mechanism and did not 
incorporate all of the requirements of the Navy spec¬ 
ifications (Section 28.2.2). The second model was 
designed to overcome certain structural defects en¬ 
countered in Model I, but, before it was built, it was 
decided to proceed with a design that would satisfy 
the Navy requirements. Therefore Model III was 
designed, built, and tested. Basically, both firing 
models were similar in that they were gas-operated, 
with blowback assist, but they differed considerably 
in the details of operation. Model III represented 
considerable improvement in design over Model I, in 
that it more nearly satisfied the Navy requirements 
and eliminated certain features of Model I that gave 
trouble in operation. These two models fired a total 
of approximately 6,000 rounds of Oerlikon and His¬ 
pano-Suiza 20-mm ammunition and also some special 
high-velocity caliber .50 ammunition. 

28 2 2 Navy Specifications for 

20-mm Automatic Aircraft Guns 

Certain broad and general specifications were sup¬ 
plied by the Bureau of Ordnance, Navy Department, 
after the project had been undertaken. These specifi¬ 
cations were supplemented by more detailed require¬ 
ments to guide the design of the third model. 329 The 
most important of them are summarized herewith, 
by quotation in part from that specification. 

1. Materials difficult to obtain or to fabricate should 
not be specified. The gun and all accessories should be 

c The work was carried out principally by Johnson Auto¬ 
matics, Inc., under Contract OEMsr-746. The kinematic 
analyses described in Section 28.2.5 were performed by the 
University of California under Contract OEMsr-1375. 


538 


CONFIDENTIAL 




JOHNSON 20-MM GUN 


539 


of a design suitable for mass production and inter¬ 
changeability of parts. 

2. The gun shall be fully automatic in operation, 
no outside power being required except for the firing 
solenoid and charger. 

3. The gun shall be capable of firing standard 20-mm 
ammunition. It shall also be capable of satisfactory 
operation with ammunition whose overall length 
is as much as }/i in. longer or shorter than stand¬ 
ard, and with ammunition so belted that the posi¬ 
tion of the bases of the rounds varies by as much 
as 14 in. 

4. With the gun mounted in an airplane wing, the 
barrel or barrel assembly shall be capable of being 
quickly and easily replaced without removing other 
parts of the gun. 

5. The feed mechanism shall be an integral part of 
the gun and energy for actuation shall be derived 
from the gun. The feed shall be interchangeable from 
right to left hand by merely exchanging components. 
The feed opening or feed tray shall be stationary with 
respect to the mounting. 

6. Cases shall be ejected downward, and links 
ejected at any anglebetweenhorizontalanddownward, 
through openings to which chutes may be attached. 

7. The maximum cross section of the gun with all 
accessories shall not exceed that of a rectangle 8 in. 
high and 10 in. wide. The overall length using the full 
length AN-M2 barrel shall not exceed 85 in. and the 
weight with all accessories shall not exceed 110 lb. 
(In conversation, Navy representatives indicated con¬ 
sideration would be given to modification of length 
and weight requirements if superior performance 
made such changes desirable.) 

8. The gun shall be capable of firing at a rate of at 
least 700 rounds per minute under service conditions. 
It shall operate satisfactorily under accelerations of 
5 g in any direction or 7 g vertically, whichever has 
the greater effect on operation. It shall be capable of 
feeding a belt of ammunition equivalent to at least 15 
rounds suspended vertically when subjected to 7g in 
the vertical direction. 

9. Except for barrel erosion, parts should be capa¬ 
ble of withstanding 5,000 firing cycles without requir¬ 
ing replacement. 

Even though the guns built on this project did not 
satisfy all the Navy’s specifications, these models 
were useful as prototypes to test the mechanism. 
When they were constructed it was recognized that 
they were subject to certain revisions to bring them 
within the specifications. 


28 2 3 Operation of Models I and III 

For a better understanding of the sections to follow, 
a detailed discussion of the sequence of operation of 
the two firing models is given herewith. 

Model I 

The lower part of Figure 1 is a vertical cross section 
through the main operating portion of Model I. Also 
shown is a top view of the breech block, breech slide 
and breech slide lever, and the principal dimensions 
of the breech slide cam path. As the bullet proceeds 
down the barrel (A), it uncovers the gas port (B), 
admitting gas at high pressure into the gas cylinder 
(C). This gas forces the piston (D) rearward, which 
in turn drives the operating shaft (E) back against 
the driving spring (F). Attached to the rear end of 
the operating shaft is the locking platform (G). A 
forward extension on the locking platform holds the 
rear edge of the lock (H) downward against a locking 
abutment on the stationary locking key (I). The for¬ 
ward edge of the lock engages a cylindrical surface on 
the bottom of the breech block (J), locking it in the 
closed position, until the locking platform is moved 
rearward sufficiently to allow the lock to rotate up¬ 
ward out of engagement with the locking key. This 
upward rotation of the lock is produced by the cham¬ 
ber pressure pushing rearward and having the locking 
abutment of such an angle (30 degrees to the vertical) 
that disengagement tends to take place. 

As the locking platform moves rearward, an ex¬ 
tension on its top surface engages a lug on the bottom 
of the striker (K) and draws it back against the 
striker spring (L). Just as the lock is released from 
the locking key, the locking platform strikes a lug at 
the lower rear end of the breech block. This impact, 
aided by residual pressure in the chamber, drives the 
breech block rearward. As the lock comes out of en¬ 
gagement with the locking key, it moves upward 
against the forward end of the locking platform, thus 
holding the breech block, striker, lock, and locking 
platform in fixed relative positions. In this manner 
they all move rearward together, being decelerated 
somewhat by the driving spring, until the breech 
block strikes the buffer spring (not shown). When 
this occurs, the moving parts are rapidly brought to 
a stop, and the energy stored in the buffer and driving 
springs drives these parts forward again. 

This forward motion continues until the forward 
end of the breech block strikes the barrel collar (N), 


CONFIDENTIAL 



540 


AUTOMATIC GUN MECHANISM 



Figure 1. Mechanism of 20-mm gun, Model I. (This figure has appeared as Figure 1 in NDRC No. Report A-454.) 


and it is brought to a sudden stop. At the same time, 
the lock is free to move downward against the locking 
key. If the sear (O) is held down by the trigger (not 
shown), the striker is free to move forward with the 
locking platform and operating shaft, under the force 
of the striker and driving springs. As the locking plat¬ 
form moves forward, its forward projection holds the 
lock in the locked position until gas pressure again 
moves the shaft and platform to the rear. 

If the trigger is released, the sear is free to move 
upward and engage the sear notch in the striker. 
Thus, when the breech block, and its associated parts 
move forward, the striker, locking platform and op¬ 
erating shaft are held to the rear of their extreme 
forward position by the sear in the sear notch, until 
the sear is pushed downward by the trigger, allowing 
these parts to move forward again. 

Ammunition for the gun is supplied from a closed- 
loop belt of disintegrating steel links, and is fed from 
the left-hand side (not shown). The feed tray is rigidly 
attached to the barrel and receiver and moves with 
them in recoil and counter-recoil, and is not inter¬ 


changeable from side to side. In these respects this 
model does not satisfy the Navy requirements, but 
since this model was constructed only to test the 
mechanism, this feature should not be considered a 
fault of the design. It was corrected in Model III. 

Power for feeding the belt is derived from the mo¬ 
tion of the locking platform. A stud (P) on the bottom 
of the locking platform fits in a curved groove in the 
feed lever. As the platform moves back and forth, the 
lever is caused to swing to the right and left. A pawl 
attached to the forward end of the lever feeds the belt 
to such a position that the round to be fed lies parallel 
to the chamber and 1% in. to the left. The feed lever 
is of such shape that the belt is fed during the closing 
stroke of the locking platform, and the lever is posi¬ 
tioned prior to feeding by the opening stroke of the 
platform. Since the details of the feed tray are not 
unlike those of other belt-fed weapons, no further de¬ 
scription of them is necessary. 

To extract the fed round from the belt and trans¬ 
late it from its position 1% in. to the left of the 
chamber to its position in the chamber, a breech slide 


CONFIDENTIAL 
































































































































JOHNSON 20-MM GUN 


541 


(R) is used. The breech slide is attached to the breech 
block by a dovetail in such a manner that it is free to 
slide to the right and left with respect to the breech 
block, and, of course, moving back and forth with the 
breech block. A cam in the top of the receiver actu¬ 
ates a lever (M), causing the breech slide to move right 
or left according to the position of the breech block. 

On the forward face of the breech slide is a T slot. 
At the left end of the breech slide are two spring- 
loaded extractor claws, giving, in effect, an extension 
of the T slot. A third extractor at the end closes the 
T slot. As the breech block and slide move forward 
into battery, the three extractor claws grip the round 
to be extracted from the belt, on three sides of the 
extractor groove. The fourth side is gripped by an 
angular face pin projecting from the forward face of 
the breech slide. Thus held securely in the breech 
slide, the round is extracted from the belt as the 
breech block moves rearward. The cam is so designed 
that lateral motion does not begin until the round is 


free of the feed tray. Part of the lateral motion takes 
place on the opening stroke &nd the remainder on the 
closing stroke. When the projectile is about halfway 
in the chamber, tl>e cam begins to return the breech 
slide to its initial position. This motion is completed 
when the breech block is about 34 in* from the bat¬ 
tery position. Thus the breech slide is moving forward 
when it engages a new round in the feed tray. 

The motion of the empty case being removed from 
the chamber is similar to that of the round being fed. 
The lateral acceleration of the breech slide is suffici¬ 
ent to throw the empty case out of the open end of 
the T slot, so the empty case is ejected forward and 
to the right of the gun at an angle of 30-45 degrees to 
the axis of the barrel. 

Model III 

Figure 2 is a vertical cross section through a portion 
of the receiver of the Model III gun. The lower view 



Figure 2. Mechanism of 20-mm gun, Model III. (This figure has appeared as Figure 2 in NDRC Report No. A-454.) 


CONFIDENTIAL 


































































































































































































542 


AUTOMATIC GUN MECHANISM 


shows the parts in the receiver as originally designed 
and built. The figure at the upper right shows the gas 
port, cylinder and piston, and that at the upper left 
shows the modifications made in the cocking mechan¬ 
ism, breech slide, and operating shaft to overcome diffi¬ 
culties encountered in operation. (See Section 28.2.4.) 

As the bullet proceeds down the barrel (A), it 
uncovers the gas port (B), admitting gas at high 
pressure into the gas cylinder (C); This gas forces the 
piston (D) rearward, which in turn drives the operat¬ 
ing shaft (E) back against the spring (F). 

As the operating shaft moves rearward Yi in., a 
raised portion (E') forces the cocking slide (G) rear¬ 
ward, which, .through a two-to-one cocking lever (H), 
cocks the striker (I) a distance of 1 in. This permits 
the sear (J) to move upward into a sear notch in the 
striker, holding it cocked against the firing spring 
(K). While the operating shaft is cocking the striker, 
its forked rear end (E") pushes the locking platform (L) 
rearward against its springs (M), until the lock (N) 
is free to rotate counterclockwise. The slope at the 
rear end of the forked operating shaft forces the lock 
to rotate. 

When the lock has rotated sufficiently, it comes 
out of engagement with a section (O') of the breech 
block (O), and the breech block is free to move rear¬ 
ward against the driving spring (P). 

At the instant the lock becomes disengaged, the 
raised portion (E r ) of the operating shaft strikes the 
forward face of the breech block. This impact, plus 
the residual chamber pressure, imparts sufficient 
energy to the breech block to drive it to the rear 
against the driving spring, until it strikes the buffer 
spring (not shown). All of the energy for feeding the 
belt and chambering the round is taken from the ki¬ 
netic energy of the breech block and the potential 
energy of the buffer and driving springs. 

Removal from the feed tray (A) of the round to be 
fed (R) is by means of a breech slide (S) similar to 
that in the Model I. The main difference fs that in the 
Model III the feed tray is above the chamber to per¬ 
mit feeding from either side, so the motion of the 
round is downward, instead of from left to right. 

The belt feed mechanism and feed tray are similar 
to the same components in the Browning caliber .50 
machine guns. The feed tray in Model III is attached 
to the mount instead of the receiver, to prevent feed 
tray motion when the gun recoils. Belt feed is inter¬ 
changeable from one side to the other, thus satisfying 
certain Navy requirements not complied with in 
Model I. 


Ignition of the round in the chamber is by means of 
the striker, the cocking of which was previously de¬ 
scribed. When the breech block closes, the lock (N) is 
rotated to the locked position, permitting the locking 
platform (L) to be forced under the rear of the lock 
by the springs (M). When the locking platform has 
moved forward a distance sufficient to assure positive 
locking, a small projection on the locking platform 
strikes a lever in the lock. Neither the projection on 
the locking platform nor the lever in the lock is shown 
in Figure 2. The lever in the lock pushes the rear end 
of the sear lever (U) upward, so its forward end draws 
the sear (J) downward, allowing the firing spring to 
push the striker forward to ignite the primer of the 
chambered round. Release of the striker cannot occur 
until the locking platform is well under the lock, thus 
preventing premature ignition with an unlocked 
breech. Manual control of firing is obtained by means 
of another sear (V), which is moved horizontally by 
an electric solenoid. When this sear is in its sear notch, 
the striker cannot move forward to strike the primer. 

In both the Model I and Model III, the round being 
fed is tightly gripped in the breech slide by the three 
extractor claws and the angular face pin from the 
time it is extracted from the belt until it enters the 
chamber. There is no opportunity for rotation during 
this period, and the round enters the chamber in the 
same angular position it had in the belt. Thus the 
guns are well suited to firing pre-engraved ammuni¬ 
tion, if the rounds are properly oriented in the belt, 
as is brought out in Section 28.4. 

28.2.4 Problems Encountered and 
Corrective Measures Applied 

It would be too much to expect that the prototype 
of any mechanism as complex as an automatic cannon 
would operate satisfactorily when first fired. Numer¬ 
ous failures occurred, and in most cases these were 
overcome by modifications in the mechanism itself. 
Some difficulties encountered would require either a 
change in the ammunition or a radical change in the 
mechanism. Insofar as possible, problems encountered 
with Model I were solved by changes in design intro¬ 
duced with Model III. 

It is quite possible that under service conditions 
other difficulties might have occurred as a result of 
conditions not encountered during testing, such as 
extremes of temperature and high acceleration of the 
gun as a whole. Space does not permit a full discus¬ 
sion of all of the problems encountered, only the most 


CONFIDENTIAL 



JOHNSON 20-MM GUN 


543 


important being discussed here. These and others are 
covered more fully in the report 126 previously referred 
to. 

“Pulled” Projectiles 

The most serious difficulty encountered, and the 
only one which was not corrected, either by improved 
design or by temporary modifications, was failure of 
the round to enter the chamber, when the guns were 
fired at cyclic rates in excess of 500 rpm. There were 
two types of failure of this nature. First, when the 
round was removed from the feed tray, the projectile 
was pulled from the case and left in the tray. Second, 
the nose of the projectile, instead of entering the 
chamber, would strike the breech face and jam the 
mechanism. 

Since these failures occurred only at high cyclic 
rates, it was evident that they resulted from dynamic 
rather than static conditions. The problem was an¬ 
alyzed by means of high-speed motion pictures, from 
which kinematic analyses were made. These analyses 
showed that the breech block undergoes very high 
accelerations (680 g at 561 rpm) when operating 
normally. Tests were made of the ammunition to 
determine what forces were required to pull the pro¬ 
jectiles from the cases. It was found that a force of 
295 lb would remove the projectile completely, and 
200 lb would loosen it badly. The acceleration force 
is of the order of 215 lb under normal conditions, so it 
might be expected that the projectiles would always 
be loosened, and under abnormal conditions would be 
pulled out completely. 

To analyze the failure of rounds to enter the cham¬ 
ber, kinematic analyses were made of the lateral ac¬ 
celerations and resulting bending moments, which 
would swing the nose of the projectile to one side or 
the other of the chamber. It was found that these 
bending moments were insufficient to cause jamming 
under normal conditions, if the projectile was tight in 
the case, but if the projectile had been loosened on 
removal from the feed tray the bending moments 
were sufficient. 

To correct these difficulties, special ammunition 
was made up which had a tighter crimp than normal. 
This functioned satisfactorily except for a few rounds 
in which the case was torn by the resulting sharp edge 
of the crimping groove. It was understood that am¬ 
munition with a stronger crimp would be available 
from the Navy Department, so further corrective 
measures were not tried. However, none of the new 


ammunition was forthcoming, so the feeding problem 
persisted for the life of the project. 

Rebound of Impacting Parts 

When two elastic pieces come together in impact, 
they tend to separate at approximately the same 
relative velocities at which they came together. This 
rebound tendency had two serious effects, “shuttling” 
and “breakage” in Model I, and reduced cyclic rate 
in Model III. 

It will be recalled from the sequence of operation 
of Model I, that when the locking platform [(G) in 
Figure 1] has moved back far enough for unlocking 
to occur, it strikes a lug at the lower rear of the breech 
block, and the block and locking platform move rear¬ 
ward together, held in fixed relative positions by the 
lock. Thus any rebound taking place after impact 
occurs by elastic deformation of the parts. Kinematic 
analysis (Section 28.2.5) showed a relative motion, 
known as “shuttling,” in which separation takes place 
by elastic deformation, to be followed by another 
impact and separation. Shuttling occurred until the 
energy of the initial rebound was absorbed by hyster¬ 
esis or plastic deformation of the affected parts. This 
frequently resulted in breakage of the lug at the lower 
rear of the breech block, and also in loss of energy 
which might have been used to increase the cyclic rate. 
This could not be corrected in Model I but was avoided 
in the design of Model III. 

Closure of the breech in Model III was completed 
by impact on the barrel collar. The locking mechan¬ 
ism was not quick enough to prevent rebound after 
this impact, with the result that the breech block 
bounced back a distance of approximately 2.7 in., in¬ 
creasing the time for the cycle by as much as 40 msec. 
Several methods were proposed to eliminate this dif¬ 
ficulty. The method finally adopted was to bring the 
breech block to a low velocity just before impact. 
This was done by holding the operating shaft to the 
rear until the breech block was about 3^8 in. from its 
battery position. Thus the breech block, just before 
going into battery, would strike the operating shaft 
and transfer most of its kinetic energy. Thus it had a 
very low velocity when striking the barrel collar and 
the lock had sufficient time to prevent rebound. 

Parts Breakage 

Structural failure of parts is always a source of an¬ 
noyance and frequently a serious danger. Fortunately 


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544 


AUTOMATIC GUN MECHANISM 


breakage of parts in either model never resulted in 
danger to personnel or other equipment, but it fre¬ 
quently did cause serious delays in the program. 

One of these failures was breakage of the breech 
block in Model I, due to “shuttling,” as previously 
discussed. The other serious failure occurred in the 
receiver plates of Model III. This was analyzed and 
found to be the result of high stresses during firing, 
applied at sections where stress concentrations and 
residual heat-treating stresses were likely to occur. 
Furthermore, the steel originally specified(SAE 4650) 
was not available in the desired sizes and shapes and 
a higher carbon steel was used. The impact strength 
of this steel was found to be quite low, which was 
probably an important factor in the failure. It was 
possible to make temporary repairs by bolting an 
auxiliary plate to the receiver plate to carry the load. 
Later, it was decided to prepare new receiver plates 
of the proper carbon content and give closer super¬ 
vision to the heat-treatment to avoid brittleness. In¬ 
sofar as possible, stress concentration was avoided in 
the redesign of the plates. 

Several other breakages occurred, but they were 
usually the result of abnormal stresses caused by 
some other failure, and would not have occurred un¬ 
der normal operation. 

Cocking Failure in Model III 

One of the first difficulties encountered in operation 
of Model III was failure of the cocking mechanism to 
operate. It was found that a pressure wave from the 
gas piston traveled down the operating shaft and 
caused the locking platform to bound away with a 
“billiard ball” effect. Thus unlocking was completed 
almost instantaneously and the breech block would 
move rearward without cocking the striker. 

This was overcome by making cocking dependent 
upon breech block travel, rather than on motion of 
the operating shaft. The changes made in the striker 
and cocking mechanism are shown in Figure 2. 

Case Ejection in Model III 

In Model I, ejection of empty cases was accom¬ 
plished by the lateral acceleration of the breech slide, 
which was sufficient to throw the empty case out of 
the T slot. It was expected that this would occur also 
with Model III, but the breech slide accelerations 
were reduced in Model III to improve chambering, 
and were insufficient to provide positive disengage¬ 


ment of the case from the T slot. The presence of the 
empty case in the T slot during the closing stroke 
would cause jamming of the mechanism. 

This difficulty was overcome by cutting away the 
T slot at its lower end and providing extraction by 
an extractor claw at the lower end of the breech slide. 
Two spring-loaded plungers, set in the face of the 
breech slide, pressed against the base of the case near 
its upper edge, tending to rotate the case in a clock¬ 
wise direction about the extractor claw. A small plat¬ 
form was bolted to the top of the operating shaft to 
support the front of the case until it was far enough 
out of the chamber that it would not strike the oper¬ 
ating shaft during ejection. These changes in the 
breech slide are shown in Figure 2. 

With the exception of “pulled” projectiles, all oper¬ 
ational difficulties encountered in testing these guns 
were overcome. It is believed that problems of feed¬ 
ing can be overcome only by a radical redesign of the 
mechanism or by a tighter crimp on the ammunition. 

28 2 5 Kinematic Analyses 

Certain difficulties in the operation of Model I in¬ 
dicated the desirability of knowing in more detail the 
actual movements of the component parts. This was 
particularly important in analyzing failures due to 
“pulled” projectiles, which were described in Section 
28.2.4. 

Kinematic analyses were made from high-speed 
motion pictures taken at 1,000 frames per second, by 
plotting displacement-time curves. Using a standard 
method of graphical differentiation, velocity and ac¬ 
celeration curves were drawn. 508 Figure 3 shows a 
typical kinematic analysis of Model I when firing at 
a cyclic rate of 561 rpm. It is noted that during the 
opening stroke, the velocity of the breech block is 
about 300 in./sec, and about 200 in./sec during the 
closing stroke. “Shuttling” of the breech block and 
locking platform appears as ripples on the velocity 
and acceleration curves in the region of +4 to +18 
msec. Ripples on the velocity curve during the clos¬ 
ing stroke are the result of transfer of kinetic energy 
to and from the breech slide. 

The high acceleration of the breech block, as a re¬ 
sult of impact from the locking platform, is shown at 
+3 msec, when it has a value of 262,000 in./sec 2 . 
Values of very high accelerations of short duration 
cannot be determined accurately, because the expo¬ 
sures at every millisecond interval do not indicate 
the actual movement during that interval but only 


CONFIDENTIAL 




JOHNSON 20-MM GUN 


545 




Figure 3. Kinematic analysis of firing cycle, 20-mm gun, Model I. (This figure has appeared as Figure 5 in NDRC 
Report No. A-454.) 


the net displacement and hence the average velocity. 
However, the values so obtained are useful in giving 
a clue to operational difficulties. 

By combining the data of Figure 3 with the draw¬ 
ings of the breech slide cam path, it was possible to 
determine the transverse accelerations of the breech 
slide. They produced bending moments tending to 
bend the projectile out of the case and the round as a 
whole out of the T slot. Curves of these bending 
moments are shown in Figure 4. These curves show 
that during the transverse movement of the breech 
slide, the moment on the complete round varies from 
about 50 to 150 lb-in., and the moment on the pro¬ 
jectile about the crimping groove varies from about 
9 to 24 lb-in. The nose of the projectile enters the 
chamber at a breech block displacement of about 7 
in. on the closing stroke, at which time the bending 
moments on the round and projectile are approxi¬ 
mately 70 and 12 lb-in., respectively. If the projectile 
were tight in the case, these moments would not be 
sufficient to prevent chambering, but it must be re¬ 
membered that the projectile was previously loosened 
by the acceleration on removal from the feed tray. 

The kinematic analysis was useful in determining 
the energy of the system. It was found that the total 


energy of the breech block and driving spring in¬ 
creased during the early part of the opening stroke, 
partially as a result of blowback. During the closing 
stroke, the energy dropped rapidly, shortly after 
leaving the buffer spring, due to the absorption of 
energy by the belt feeding mechanism, thus giving a 
slow closing stroke. If energy for belt feeding had 
been taken near the end of the closing stroke, instead 
of at the beginning, a considerably higher cyclic rate 
would have been obtained. 

There was insufficient time to make complete kine¬ 
matic analyses of Model III. However, the an¬ 
alyses did show a very high acceleration (900,000 in./ 
sec 2 ) resulting from impact of the operating shaft. 
These curves were similar to those for Model I, ex¬ 
cept for the presence of rebound at the end of the 
closing stroke, and the absence of “shuttling.” 

The kinematic analyses were probably the most 
important factor in locating and eliminating sources 
of trouble in operation of the guns. 

28.2.6 Prediction of Firing Rate 

During the design period of Model III, it was felt 
desirable to predict the firing rate, in order that spring 


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546 


AUTOMATIC GUN MECHANISM 


constants and other factors might be given optimum 
values before completion of the design. Unfortunate¬ 
ly the time available was too short to incorporate 
the results of the calculations into the design, or to 
make a complete calculation of the final design. 
Hence the calculations shown herein apply only to a 
hypothetical gun having the same general character¬ 
istics as the Model III gun. 

Briefly the method of calculation consisted of set¬ 
ting up a differential equation of motion for each part 
of the cycle, determining the time for that part, and 
taking the sum of the times as the total for the cycle. 
For some parts of the cycle the time was known or 
estimated, some could be calculated directly, and 
others had to be calculated by a step-by-step method. 


The cycle was considered to start at the instant the 
sear released the striker. The results of the compu¬ 
tations are summarized in Table 1. The bases on 
which they were made are indicated in the following 
paragraphs, which are numbered to correspond to the 
successive time intervals. 

1. The striker had simple harmonic motion, being 
a simple spring-mass system for which the mass 
weighed 1.02 lb and the spring had a constant of 30 
lb/in. The natural frequency co of the system was 
evaluated by equation (1) 



in which k is the elastic constant of the spring and m 




OPENING-CLOSING-— 

Figure 4. Bending moment on round due to breech slide acceleration, 20-mm gun, Model I. (This figure has appeared 
as Figure 6 in NDIiC Report No. A-454.) 


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JOHNSON 20-MM GUN 


547 


Table 1 . Calculated time for each interval in the firing 
cycle of 20-mm automatic cannon. 


Inter¬ 

val* 

Event 

Time 

(sec) 

<1 — to 

Striker motion 

0.0125 

u - h 

Projectile to pass port 

0.0016 

u - U 

Operating shaft motion 

0.0015 

1 4 - U 

Primary extraction 

0.0002 

<5 — u 

Impact 

0.0002 

U> — ti 

Blowback 

0.0067 

<7 — U 

Motion of breech block from the end of 



blowback until engagement with buffer 

0.0134 

ti — <7 

Compression of buffer 

0.0074 

<9 ~ U 

Decompression of buffer 

0.0074 

tio — 1 9 

Buffer to battery 

0.0380 

hi — ho 

Locking platform motion 

0.0096 


Total 

0.0985 


Equivalent firing rate: 609 rpm 



* The cycle is considered to start at the instant the sear releases the 
striker. 


is its mass. The time for the first interval of the cycle 
is then given by equation ( 2 ), 

ti — to = — cos -1 — ( 2 ) 

00 Xo 

in which x 0 and X\ are the compressions of the spring 
before release and after striking the primer. 

2. Interior ballistic data for the ammunition fired 
showed that the time t 2 — t\ for the projectile to move 
11 in. after the primer was struck was 0.00159 sec. 

3. After the gas port had been uncovered, high 
pressure gases flowed into the cylinder and drove the 
piston and operating shaft rearward. Equation (3) is 
the general equation of motion. 

mx + kx = F, (3) 

where F is the force on the piston and x its displace¬ 
ment. This force is a function of the pressure in the 
barrel (which varies with time) and of the volume in 
the cylinder (which varies with x). Inasmuch as it is 
impractical to use an analytical function of F, the 
motion was solved by a step-by-step process, using 
time intervals of 0.0001 sec to find the time required 
to move 0.5 in. 

4. During the next J^-in. movement of the oper¬ 
ating shaft, it cammed the lock out of engagement 
with the breech block and imparted some initial 
energy to the breech block. This “primary extraction” 
process was most easily solved by an energy balance, 
making assumptions regarding friction and taking 
into account the effect of blowback. 

5. The laws of impulse and momentum were used 


to estimate the velocity of the breech block after im¬ 
pact. 

6 . During the early part of the opening stroke of 
the breech block, the motion was aided by blowback 
and retarded somewhat by friction. The equation of 
motion is similar to equation (3), but, of course, has 
different constants. Since the blowback force was not 
a constant or a simple function of x or t, a step-by- 
step process was used. Friction was estimated at 50 
lb and the blowback pressure p at any instant was 
calculated 48 by equation (4) 


in which p 0 is the barrel pressure when the projectile 
leaves the muzzle, a is the ratio of the muzzle veloc¬ 
ity to the barrel length, 7 is the ratio of the specific 
heats of the powder gas, and the time t' is measured 
from the instant the bullet leaves the muzzle. 

7. The motion of the breech block from the end of 
blowback until it struck the buffer spring was similar 
to that during blowback except the blowback force 
was zero. 

8 . The required constants of a buffer spring to be 
compressed 2 in. were calculated, and the new value 
of k substituted in equation (3) for calculation of the 
period of buffer spring compression. 

9. It was assumed that decompression of the buffer 
spring required the same time as compression. 

10 . The time for the closing stroke of the breech 
block was calculated in two different ways; first, by 
application of the principles of a spring-mass system 
taking into account the energy required for belt feed¬ 
ing, and, second, by assuming that the closing stroke 
required about 65 per cent longer than the opening 
stroke. The value obtained by the latter method was 
used. 

11 . After the breech block went into battery the 
locking platform had to move under the lock. Again 
the process involved a simple spring-mass system. 

The total calculated time for the cycle was 0.0985 
sec, so that the calculated firing rate was 609 rounds 
per minute. 

Two possible ways of increasing the firing rate with¬ 
out affecting breech block motion are by shortening 
the time for the striker to move and by using an 
instantaneous locking device. The first of these was 
actually done to correct another fault (cocking diffi¬ 
culties—see Section 28.2.4), and the time for striker 
motion reduced to 0.005 sec, which would increase the 
calculated firing rate to 659 rpm. The use of an in- 


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548 


AUTOMATIC GUN MECHANISM 


stantaneous lock such as a toggle link would reduce 
the time for the cycle by a further 0.0096 sec, raising 
the calculated firing rate to 737 rpm. 

28 2 7 Conclusions 

As a result of the experimental work and theoretical 
analyses performed on this 20-mm gun project, cer¬ 
tain conclusions may be drawn. 

1. Both models that were built and fired per¬ 
formed satisfactorily within limits. 

2. All of the problems encountered, with the excep¬ 
tion of “pulled” projectiles, were overcome by changes 
in the gun mechanism itself. 

3. Neither of the guns built completely satisfied 
the Navy requirements with respect to weight and 
length. 

4. The standard Navy ammunition used is not 
satisfactory for this type of mechanism, without 
modifying the mechanism to reduce acceleration 
forces and bending moments. 

5. The kinematic analyses were of great importance 
in determining causes of failure and in correcting 
them. 

6. Both models had remarkable belt-pulling abil¬ 
ity. 

7. Impacts resulting in “shuttling” or rebound 
cause parts failure and low cyclic rates. 

8. More information on the action of gas cylinders 
and blowback is needed. 

9. During the design stages, the results of theo¬ 
retical kinematic analyses (calculation of firing rate, 
etc.) would be helpful in choosing correct springs and 
weights. 

10. Quick-acting locking devices are useful to in¬ 
crease cyclic rate. 

11. High breech block weights require more energy 
for the same cyclic rate. 

12. Long travel for parts reduces cyclic rates if vel¬ 
ocities and accelerations are to be kept to a minimum. 

13. Maximum blowback power is obtained with 
early breech opening. This is limited, however, by 
the ability of the case to resist rupture with high in¬ 
ternal pressure. 

14. The gun mechanism is adaptable to pre-engraved 
projectiles. 

28 2 8 Recommendations 

In order to make the experience of this project of 
most effective use to future designers, there are pre¬ 


sented herewith certain recommendations which the 
author believes will minimize difficulties if they are 
adopted. Some of these recommendations are very 
broad while others are specific, some relate to design 
and others to further lines of investigation, and some 
concern the gun alone while others concern the am¬ 
munition in relation to the gun. 

1. The rounds should be moved with accelerations 
as low as possible to avoid inertia forces and couples 
that tend to loosen projectiles in the case and make 
chambering difficult. This applies to belt feeding, re¬ 
moval from belt and translation to the axis of the 
chamber. This might be accomplished by the use of 
buffer springs to limit maximum accelerations. 

2 . To take maximum advantage of blowback pow¬ 
er, the breech should be opened as early as practic¬ 
able. There are some limitations to this expedient 
however. Extraction of energy for blowback may re¬ 
sult in loss of muzzle energy, and early extraction 
subjects the empty case to high internal pressures 
which might rupture it and make extraction difficult. 

3. The ammunition should be designed for the gun. 
This design should include sufficient crimping to pre¬ 
vent loosened projectiles, and sufficient wall strength 
of the case to permit early extraction. This recom¬ 
mendation is closely allied to recommendations 
1 and 2 . 

4. For maximum cyclic rate, the breech block open¬ 
ing and closing strokes should be at as high a velocity 
as practicable. For limited total energy of the block, 
the weight should be low and the driving spring only 
stiff enough to assure breech block closure under the 
worst conditions. 

5. Energy for auxiliary functions, such as cocking 
and belt feeding, should be extracted as near the end 
of the breech closing stroke as possible, to maintain a 
high kinetic energy in the breech block. This has the 
added advantage of reducing the breech block veloc¬ 
ity as it goes into battery and helps eliminate re¬ 
bound. 

6 . For better analysis of gas-operated weapons, an 
analytical study should be made of the action of gas 
ports and pistons. 

7. The time required for events in the cycle that 
occur between the time the breech block closes and 
the striker hits the primer should be as short as pos¬ 
sible, to get high cyclic rates. Instantaneous locking 
devices and short stroke strike rs will shorten this time 
considerably. 

8 . The weights of the operating shaft and the breech 
block should be so proportioned that the maximum 


CONFIDENTIAL 





AUTOMATIC MECHANISM FOR HYPERVELOCITY AMMUNITION 


549 


amount of energy is transferred from the shaft to the 
breech block at the instant of impact. 

9. Rebound of the breech block, when going into 
battery, can be eliminated by having it impact against 
some mass, such as the operating shaft, thus losing 
kinetic energy before the impact with the barrel col¬ 
lar or other stationary mass. 

10. Careful dynamic studies should be made dur¬ 
ing design, to avoid faults that would result in low 
cyclic rate or failure to operate under conditions of 
high acceleration. 

11. To keep the gun as short and light as possible, 
the length and travel of the breech block should be 
kept short. This will also raise the cyclic rate. 

12. Insofar as possible, locking stresses should 
be limited to the forward part of the receiver, or 
preferably kept free of the receiver. In this way 
the weight of the receiver can be materially re¬ 
duced. 

13. If locking devices, driving and buffer springs, 
and the weight distribution of the breech block are 
symmetrical about the axis of the chamber, friction 
forces and lateral stresses could be minimized and the 
weights of parts reduced. 

14. The charger and trigger solenoid should be 
mounted inside the receiver, so that right- and left- 
hand interchangeability is not required. This would 
reduce the number of openings required in the re¬ 
ceiver and the stress concentrations resulting there¬ 
from. 

15. Snap catches, etc., that have to act immediately 
should be avoided, because they tend to engage on 
very small surfaces and rapid wear occurs. 

16. Buffer springs between impacting surfaces 
might be used to reduce impact stresses and high 
accelerations resulting from impact. 

17. Spring-actuated parts should have springs of 
sufficient strength to prevent undesired motion of 
these parts under high “g” conditions. 

18. Ejection ports should be sufficiently large and 
free of obstructions to prevent jamming of empty 
cases and links. It must be remembered that com¬ 
plete rounds must frequently be ejected in case of 
misfire. 

19. Large flat plates for the receiver or other parts 
should be avoided, because they are difficult to heat- 
treat without distortion. 

20. The parts should be designed so they can be 
easily machined from bar, tube, or plate stock with¬ 
out removal of too much material. If necessary, forg¬ 
ings or castings could be used, but the use of stock 


material is preferable from the procurement stand¬ 
point. 

21. Parts requiring high fensile strength should not 
be subjected to impacts and should be designed to 
avoid stress concentration and residual heat-treat¬ 
ment stresses. Likewise parts subjected to impacts 
should not be required to carry high stresses. This is 
necessary to avoid the use of special alloys having 
both high tensile and high impact strengths. 

22. Consideration should be given to modification 
of the power linkage by the use of multiple-drive 
units, one gas cylinder being used to produce fore- 
and-aft movements and another to produce move¬ 
ments of the belt and other parts moving at a right 
angle to the axis of the barrel. In this way a higher 
cyclic rate could be obtained without putting undue 
stress on the moving parts. d 

28 3 PROBLEMS IN DESIGNING 
AUTOMATIC MECHANISM FOR 
HYPERVELOCITY AMMUNITION 

28 3 1 Introduction 

Higher velocity of a given projectile may be ob¬ 
tained by increasing the mean net force acting on the 
base of the projectile, or by increasing the distance 
over which that force works. The distance over which 
the force acts can readily be increased by lengthening 
the barrel of the gun. The mean net force acting 
depends upon the height and shape of the pressure- 
travel curve, and can be increased by increasing the 
maximum gas pressure, decreasing retarding forces 
(friction), or by maintaining a larger proportion of 
the curve at or near its maximum value. 

There are objections to the use of any of these 
methods of increasing muzzle velocity, so a com¬ 
promise is usually made of the various factors. Longer 
barrels result in heavier guns that require more power 
for the elevating and traversing mechanism. Increased 
maximum gas pressures mean that the walls of the 
barrels must be thicker (and heavier) or made of 
higher strength materials. Little can be gained by 
reducing friction, as this force is only a small fraction 
of the total propelling force. In order to maintain a 


d This recommendation is taken from a postscript written 
for NDRC Report A-454 126 by Mr. J. A. TenBrook, formerly 
Head of the Engineering and Development Branch, Division 1, 
NDRC. In that position he had supervision of the contract 
under which the Johnson 20-mm gun was being developed. 
(Editor’s note.) 


CONFIDENTIAL 





550 


AUTOMATIC GUN MECHANISM 


larger proportion of the pressure curve at or near its 
maximum value, that is, to increase the ratio 

mean effective pressure 
maximum pressure 

a larger quantity of slower burning powder is re¬ 
quired. This increases heat transfer to the barrel, 
gives higher muzzle pressures, and increases muzzle 
flash. An increased powder charge usually requires a 
larger cartridge case. This may be obtained by in¬ 
creasing its length, diameter, or both. 

The tactical disadvantages of the methods of ob¬ 
taining higher muzzle velocities have been mentioned. 
These methods also introduce serious problems in the 
design of the gun mechanism itself. 

The remarks to follow relate only to weapons using 
normal weight projectiles, 6 in which high velocities are 
obtained by increased pressure, longer barrel and in¬ 
creased powder charge. They are not concerned with 
the use of “‘half-weight/’ sabot or deformable pro¬ 
jectiles. 


Recoil 


The maximum velocity of the recoiling parts in free 
recoil V s is given 509 by equation (5), 


V f 


MV + 4700C 
W 


(5) 


in which M is the weight of the projectile, V is the 
muzzle velocity of the projectile, C is the weight of 
powder, 4700 is the assumed mass-mean velocity of 
the powder gas, and W is the weight of recoiling 
parts. If the gun is to be operated as a hypervelocity 
gun, V and C would be increased and the mass-mean 
velocity of the gas would probably be increased some¬ 
what due to higher muzzle pressures and tempera¬ 
tures. The weight of the gun would probably be some¬ 
what greater to give greater strength or longer travel, 
but in general the numerator of equation (5) would 
increase faster than the denominator. Therefore the 
velocity of free recoil would be higher at hyperveloc¬ 
ities than at normal velocities. This would necessitate 
a redesign of the recoil mechanism to absorb greater 
energy, permit longer travel or exert greater retarda¬ 
tion. In the case of recoil-operated guns, such as the 
Browning machine gun, hypervelocity ammunition 
and the greater recoil velocity (with normal weight 


e That is, projectiles for which the ratio of the weight M to 
the cube of the diameter d is approximately 0.5 lb/in. 3 . 


projectiles) would probably give increased cyclic rate. 

28 3 3 Barrel Design 

Needless to say, the use of higher chamber pressures 
to secure higher velocities necessitates heavier walls 
or higher strength materials. Dynamic stresses re¬ 
sulting from increased recoil accelerations would have 
to be given consideration, as would bending stresses 
due to barrel whip if the length were increased. Barrel 
temperatures would be somewhat higher (see Chap¬ 
ter 5), so the effect of temperature on strength would 
be important and provisions for adequate cooling 
might be required. 

28 3 4 Breech Lock or Breech Ring 

Inasmuch as high muzzle velocities are usually ob¬ 
tained by the use of higher chamber pressures the 
strength of the breech lock must be increased. This 
condition is further aggravated by the fact that the 
cartridge case is of larger diameter giving an addi¬ 
tional increase of total force on the breech block. If a 
sliding breech block is used, these remarks apply to 
the breech ring. 

28 3 5 Breech Block 

Sliding breech blocks frequently have an insert on 
the forward face to facilitate assembly of the striker 
mechanism. High chamber pressures and setback of 
the cartridge case occasionally produce enough de¬ 
formation of the breech block and insert to interfere 
with opening of the breech. Therefore consideration 
must be given to the strength of the breech block to 
avoid this condition. 

To accommodate the larger powder charges re¬ 
quired for hypervelocity, the cartridge cases are 
usually increased in length and diameter. Therefore 
to handle the longer round, recoiling breech blocks 
must travel a greater distance and the receiver and 
mechanism contained therein must be modified ac¬ 
cordingly. 

28,3,6 Feed Mechanism 

Since the total weight of the complete hyperveloci¬ 
ty round is somewhat greater than that for normal 
velocity, the feed mechanism must be somewhat 
stronger, and the mechanism will require larger quan¬ 
tities of powder for satisfactory operation. 


CONFIDENTIAL 






LOADING AND INDEXING MECHANISMS 


551 


28 4 LOADING AND INDEXING 
MECHANISMS FOR PRE¬ 
ENGRAVED PROJECTILES 

28 4 1 Introduction 

Pre-engraved projectiles in chromium-plated bar¬ 
rels are useful in prolonging barrel life, as is shown in 
Chapter 31. In order for such projectiles to be prac¬ 
tical, there must be some means of indexing them 
positively, to make certain that the splines or teeth 
on the projectile will engage the rifling of the barrel 
during chambering of the round. If proper orienta¬ 
tion is not obtained, the splines will jam on the lands 
and result in excessive powder pressure and rapid 
wear at the origin of rifling. 

Tests 120 made on existing loading mechanisms 
showed how much angular rotation of the projectile 
took place in moving from the belt or clip or other 
ammunition supply until it was seated in the cham¬ 
ber. By pointing the lands and splines with matching 
30-degree slopes a certain amount of misalignment 
could be tolerated, but it would be preferable to have 
accurate alignment. 

The Browning caliber .50 machine gun showed a 
case rotation of ± 20 degrees, which would be satis¬ 
factory for a four-land barrel. Large and irregular 
rotation of the cartridge case was found in the 37-mm 
AA gun, M1A2 and in the 40-mm AA gun, Ml (Bo- 
fors). About half the shells were inserted unfavorably 
in hand loading and in power loading a 90-mm gun, 
Ml, at 22 rounds per minute. These tests indicated 
that changes in the projectile, origin of rifling, and 
feeding mechanism were required to chamber the pre¬ 
engraved round satisfactorily. 


The following guns and ammunition were modified 
to fire pre-engraved ammunition: ( 1 ) Caliber .50 
Browning machine gun, ( 2 ) 40-mm Bofors gun, and 

(3) 20-mm Johnson gun, Model III. 

28 4 2 Browning Caliber .50 Machine 
Gun 

Several changes were made in the gun and ammu¬ 
nition to make possible the chambering of pre-en¬ 
graved ammunition, as follows: 

(1) Four rifling grooves in the barrel instead of 
eight. 

( 2 ) Groove depth 0.020 in. instead of 0.005 in. 

(3) Pre-engraved steel bullet with four integral 
splines 0.020 in. high, instead of soft jacket. 

(4) Splines on bullet pointed with 30° slope. 

(5) Lands at origin of rifling pointed with 30° 
slope. 

( 6 ) Indexing slot on base of cartridge case. 

The splines on the projectile were properly ori¬ 
ented with respect to the indexing slot, which in turn 
was given a definite angular position in the belt. 
These two steps were easily taken by modification of 
the cartridge and belt loading mechanisms. Thus the 
projectiles were so aligned in the belt that, if no rota¬ 
tion took place in going from the belt to the chamber, 
all rounds would occupy the same angular position in 
the chamber. Only four ratchet notches are milled at 
the rear of the barrel, so the barrel would have to be 
in one of the four desirable angular positions. Previ¬ 
ous tests had shown a maximum angular displace¬ 
ment of the round of only + 20 degrees in going from 
the belt to the chamber. This is well within the allow- 



Figure 5. Caliber .50 pre-engraved projectile for test of automatic fire in Browning machine gun. (This figure has 
appeared as Figure 30 in NDRC Report No. A-448.) 


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552 


AUTOMATIC GUN MECHANISM 




,5 30 " +. 00 ! 

DIA - 000 " 

AFTER PLATING 


ENLARGED SECTION THROUGH 
BORE SHOWING RIFLING 

Figure 6 . Chamber details of caliber .50 machine gun barrel modified to permit firing the pre-engraved projectile 
shown in Figure 5. (This figure has appeared as Figure 34 in NDRC Report No. A-448.) 


able limits to prevent “hang-ups” in a four-groove 
barrel, with pointed lands and splines. 

In firing 600 rounds of pre-engraved ammunition 
in this gun, there were no malfunctions in entering 
the rifling and no hang-ups at the origin of rifling, in¬ 
dicating that this method of indexing is satisfactory 
at rates of 500-550 rpm. 

The pre-engraved projectile is shown in Figure 5, 
and the chamber details of the modified barrel in 
Figure 6. 

28 4 3 Johnson 20-mm Automatic 
Aircraft Guns 

The Model I and Model III guns, described in Sec¬ 
tion 28.2.3, were adapted to fire high-velocity caliber 
.50 ammunition. These guns employ a novel feeding 
system whereby the round is positively held in the 
breech slide from the time it is drawn from the belt 
until it is well in the chamber, with no opportunity 
for rotation about its axis. Thus, if the round is prop¬ 
erly indexed in the belt, it will be properly oriented 
in the barrel. 


The tests in the Model I gun were simply to deter¬ 
mine whether or not such orientation was actually 
obtained. They consisted of putting a scratch on each 
case and loading the rounds in the belt so the scratches 
Avere at the 12 o’clock position. A corresponding 
groove was cut in the 12 o’clock position in the cham¬ 
ber, so the high chamber pressure of firing would 
leave an indicating mark on the case. This could be 
compared Avith the scratch previously made to indi¬ 
cate the amount of rotation, if any. The ammunition 
consisted of a caliber .50 ball bullet, M2, loaded into 
a 20-mm case necked down to fit the caliber .50 pro¬ 
jectile. A charge of approximately 476 grains of dou¬ 
ble-base powder was used to give a pressure of 56,000 
to 58,000 psi and a velocity of 3,600 to 3,700 fps. In 
tests at 560 to 570 rpm, it was found that very little, 
if any, rotational displacement of the rounds took 
place, indicating that this mechanism would be suit¬ 
able for pre-engraved ammunition. 

The Model III gun was fired with caliber .50 pre- 
engraved high-velocity ammunition. This was similar 
to that used in Model I, except that pre-engraved 
projectiles with pointed lands Avere used as had been 


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LOADING AND INDEXING MECHANISMS 


553 



Figure 7. Complete rounds of 37-mm pre-engraved ammunition assembled in clip ready for loading into 37-mm gun. 
(This figure has appeared as Figure 38 in NDRC Report No. A-448.) 


done in the Browning caliber .50 barrel. In the 50 pre¬ 
engraved rounds fired, the rotation of the assembled 
rounds was negligible. In fact, it was found that even 
with random indexing in the belt, the pointed lands 
and splines were sufficient to give correct orientation 
in the chamber. 

28 4 4 Modified Bofors 40-mm Mechanism 

This gun was modified to fire 37-mm pre-engraved 
ammunition by using six grooves of 0.030-in. depth, 
with the lands pointed on a 30-degree slope at the 
origin of rifling and by changing the feeding mechan¬ 
ism to prevent rotation of the previously indexed 
round. The projectile had six corresponding splines 
pointed with a 30-degree slope to match the lands. 


They were aligned in the case so that proper indexing 
of the case would automatically index the round. The 
case was the standard 40-mm case, necked to fit the 
37-mm pre-engraved projectile. A guiding groove Avas 
cut in the base to guide the case on a vertical aligning 
track. An extension of this groove was cut in the edge 
of the flange to guide the case along a horizontal track 
in the tray. The complete rounds f assembled in the 
clip are shown in Figure 7. 

The feed mechanism was modified by insertion of 
guiding tracks to maintain alignment of the rounds. 
A vertical track was mounted on the base guide to 
engage the slot milled in the base of the case, as it 


f A single round and separate projectile are pictured in 
Chapter 31 (Figure 18). 


CONFIDENTIAL 




554 


AUTOMATIC GUN MECHANISM 


FOR AUTOMATIC FIRING OF PRE-ENGRAVED PROJECTILES IN 40’MM MOUNT 


GUIDE 

TRACK 



FIRST STAGE “GUN LOADED 



FIRED CASE 


SECOND STAGE-CASE BEING EJECTED 

Figure 8A. First two stages in the operation of the automatic loading mechanism for 37-mm pre-engraved ammuni 
tion. (This figure has appeared as Figure 41 in NDRC Report No. A-448.) 


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LOADING AND INDEXING MECHANISMS 


555 


moved downward. A similar track was mounted on 
the feed tray and extended in sections on the breech 
ring and breech block to engage the groove in the 
edge of the flange. Thus while the round was moving 
vertically to the feed tray it was guided by one track 


and groove, and by another track and groove while 
moving horizontally. A sbring-loaded guide latch 
acted as an extension of the vertical track to align the 
case during the transition from the vertical to the 
horizontal track. While the round was being rammed 



THIRD STAGE-CARTRIDGE ON LOADING TRAY 



FLEXIBLE GUIDE PREVENTS 
CASE FLANGE FROM LEAVING 
GUIDE TRACK 


FOURTH STAGE— CARTRIDGE BEING RAMMED 

Figure 8B. Last twcTstages in the operation of the automatic loading mechanism for 37-mm pre-engraved ammunition. 
(This figure has appeared as Figure 41 in NDRC Report No. A-448.) 


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556 


AUTOMATIC GUN MECHANISM 


by the feed rollers, a flexible guide above the case pre¬ 
vented the case flange from leaving the horizontal 
guide track. Figure 8 shows four stages in the firing, 
ejection of an empty case, and chambering of a new 
round. 

This work was not completed under NDRC aus¬ 


pices. It was continued during 1946 on an Army con¬ 
tract® with the plan that the mechanism will be tested 
for functioning of pre-engraved ammunition under 
automatic firing conditions. 


g Contract W-36-034-7380 with the Franklin Institute. 


CONFIDENTIAL 






Chapter 29 

SABOT-PROJECTILES 

By J. S. Burlew h 


i 


29 i THE SABOT PRINCIPLE 

291,1 Introduction 

T he development of sabot-projectiles was one of 
several projects undertaken by Division 1 in its 
endeavor to attain higher useful muzzle velocities 
than were attainable with conventional projectiles 
fired from conventional guns. A sabot-projectile may 
be fired from a standard gun interchangeably with 
standard rounds without alteration to the gun or 
change in the normal powder pressure. It was this 
feature of sabot-projectiles that made their develop¬ 
ment seem especially worth while at the beginning of 
this country’s participation in World War II. The use 
of heavier armor by the Germans, as has been brought 
out in Chapter 1, had created a need for higher veloc¬ 
ity armor-piercing projectiles, in order to increase the 
penetration at battle ranges of existing guns. 

The hope that this need might be met most quickly 
by firing sabot-projectiles from existing guns led 
NDRC to focus attention on their development. Al¬ 
though these projectiles were not developed in time 
to be used by American forces during the war, the 
latest tests indicated that a 90-mm sabot-projectile 
could be made so that its accuracy would be as great 
as that of the standard projectile, while at the same 
time it would have greater armor penetration because 
of its higher velocity. 

29,1,2 General Description 

A sabot-projectile is one that separates into two or 
more parts after leaving the muzzle of the gun. One 

a As originally planned, this chapter was to have been a 
critical analysis of the problem of designing sabot-projectiles, 
illustrated by examples drawn from the experience of Division 
1 contractors. It was not possible to find an author who had 
the time to do this, and therefore the Editor has supplied 
merely a resume of what Division 1 did in this field, with 
references to the formal NDRC reports in which further de¬ 
tails may be found. He is indebted to Mr. J. McG. Millar 
(formerly Assistant Physicist, Geophysical Laboratory, Car¬ 
negie Institution of Washington) and to Mr. Edmund Bain- 
bridge (formerly Engineer, NDRC Engineering and Transi¬ 
tion Office) for helpful suggestions concerning the material in 
Section 29.1 and to Mr. Millar for a draft of Section 29.2.2. 

b Technical Aide, Division 1, NDRC. (Present address: 
Geophysical Laboratory, Carnegie Institution of Washington.) 


or more parts comprising the sabot are discarded, 
while the projectile proper (referred to as the sub¬ 
caliber projectile) proceeds along its trajectory. 0 Fig¬ 
ure 1 is an exploded view of the individual parts of a 
sabot-projectile of the same general design as that 
described in Section 29.3.4. 

The muzzle velocity of a sabot-projectile is greater 
than that of the projectile of normal weight for the 
same caliber gun, when the powder charge is the same, 
because of the reduced mass of the projectile. The 
mere attainment of increased muzzle velocity is not 
particularly difficult, but by itself it is of no practical 
value. In order for the hypervelocity sabot-projectile 
to be useful, it must have a good form factor and satis¬ 
factory stability, so that it retains a higher velocity 
all the way to the target. Therefore the first step in 
the program for the development of sabot-projectiles 
by Division 1 was a study (described in Section 29.2.2) 
to determine the design limitations for the subcali¬ 
ber projectile in order that satisfactory stability 
might be attained when it is fired from existing guns 
having standard twist of rifling. 

29,1,3 Advantages and Disadvantages of 
Sabot-Projectiles 190 

Value for Armor Penetration 

The application of the sabots with which Division 
1 was concerned for the most part was intended to 
increase the armor penetration obtainable with guns 
already in service use. 100 Even though a considerable 
portion of the energy of a sabot-projectile is lost be- 

c In many reports dealing with sabot-projectiles the caliber 
of the subcaliber projectile has been included in the designa¬ 
tion of the caliber of the projectile, as for instance: “11-20 
mm,” “75-57 mm,” and “105-75 mm.” No standard designa¬ 
tion has been established. In the present chapter such desig¬ 
nations have not been used, for in the light of the later develop¬ 
ment work on sabot-projectiles it does not seem necessary to 
place special emphasis on the diameter of the subcaliber pro¬ 
jectile. It is no more fundamental a property than some other 
characteristics, such as the weight, which are not included in 
the designation. Furthermore, to refer to a sabot-projectile 
simply as a 90-mm one, instead of as a “90-56 mm” one, for 
example, emphasizes the important point that as far as use is 
concerned the sabot-projectile is interchangeable with other 
rounds for the same gun. 


CONFIDENTIAL 


557 




558 


SABOT-PROJECTILES 





BEARING 

BANO 


TRACER POCKET 
AND SHEATH 



CORE 



WINDSHIELD 


Figure 1. Exploded view showing the separation of the various parts of a sabot-projectile, based on the design of the 
90-mm projectile shown in Figures 7 and 8. 


cause of the high ratio of charge to projectile weight 37 
and because of the nonpenetrating parts of the sabot, 
the remaining energy is concentrated over a much 
smaller area of the armor plate. 

A steel subcaliber projectile does not show a large 
advantage for armor penetration. 14 42 By the proper 
choice of diameter of the subcaliber projectile in a gun 
with favorable twist, an increase in armor penetration 
of about 20 per cent over the standard projectile is 
obtainable, provided that shatter (Section 9.2) does 
not decrease its effectiveness. 

As is pointed out in Section 9.2, the properties of 
good tungsten carbide are such that it is superior to 
steel for armor penetration, especially at hyperve¬ 
locities. The optimum size core depends on a multi¬ 
tude of factors, such as the best length of core relative 
to the diameter, the weight of the sabot, the twist of 
rifling of the gun, and the range at which the maxi¬ 
mum amount of armor penetration is desired. 14 Some 
of these factors are discussed in Section 33.2.4 


In addition, a sabot may make possible the use of 
a lighter gun for firing a given projectile at a given 
velocity. This advantage would of course be especially 
important for aircraft armament or for airborne 
equipments A complex calculation 205 - 551 involving 
ballistics, strengths of materials, and carriage weight, 
among other factors, would be required for its evalu¬ 
ation for any particular use. 

Effect on Erosion 

So far there have not been enough sabot-projec¬ 
tiles tested to determine the rate of erosion of a gun 
firing such ammunition; but analysis 48 of its ballistics 
indicates that the erosion might be less than that of 
a gun firing conventional ammunition. Another pos¬ 
sibility is that of shooting fin-stabilized projectiles 
from a smooth bore gun (Section 29.2.2). The latter 
feature might simplify construction of the gun and 
decrease the erosion. 


Other Advantages 

There are several other uses of sabot-projectiles 
which have been considered by Division 1 in a pre¬ 
liminary fashion. One of these is to effect an increase 
in the maximum range of the projectile over that of 
the standard projectile for a given gun. 42 Thus it has 
been computed that shooting a 10-in. projectile from 
a 16-in. gun would increase the range of this gun at 
least 12,000 yd, that is, from 49,000 to 61,000 yd. 

Another advantage is the possibility of shortening 
times of flight of medium-caliber projectiles at long 
ranges. By the use of fin stabilization it has been 
computed that times of flight at ranges of the order of 
30,000 yd could be reduced by 30 to 40 per cent. 
Although the projectile proposed would probably be 
effective for direct hits, there might be no advantage 
in shells with proximity or time fuzes because of the 
reduced high explosive charge. 41 In this connection 
see Section 33.2.3. 


Disadvantages 

There are of course disadvantages to a sabot-pro¬ 
jectile. The principal one is the danger to friendly 
troops. If a sabot is a centrifugal type (Section 29.2.3), 
the parts fly off in a cone with an apex angle deter¬ 
mined in large part by the twist of rifling. It is approx¬ 
imately 14 degrees for a twist of 1:25. Although their 
range is short, there is still a distinct possibility that 
the fragments will hit friendly troops. The danger 
from the axial type (Section 29.2.3) is less since all 
parts travel in a line parallel to the trajectory, but on 
hitting the ground ricochets at any angle are possible. 
However, for use in tanks and airplanes (especially 
jet-propelled ones) these objections might be minor. 

d This objective was achieved to a limited extent in the de¬ 
sign of a sabot for firing the 57-mm projectile, M86, from the 
105-mm howitzer M3, as described in Section 29.3.5. The 
57-mm gun, Ml, mounted on a field carriage weighs 2,700 lb, 
whereas the 105-mm howitzer, M3, on its mount weighs only 
2,495 lb. 


CONFIDENTIAL 











DESIGN OF SABOT-PROJECTILES 


559 


In comparison with a skirted projectile fired from 
a tapered bore gun, it is brought out in Section 33.2 
that a sabot-projectile has the disadvantage of losing 
some of its energy in the discarded sabot parts. This 
disadvantage is counterbalanced by the fact that a 
sabot-projectile can be fired from a standard gun 
interchangeably with regular ammunition. Hence 
each type of subcaliber projectile is likely to have its 
own field of usefulness in the future. 

29 2 DESIGN OF SABOT-PROJECTILES 
29 2 1 General Requirements 

A complete sabot-projectile must combine a sub¬ 
caliber projectile suited to its purpose (either armor 
penetration or high-explosive use) with a sabot that 


A number of patents 212 on sabots for subcaliber 
projectiles have been granted. Although detailed 
specifications are in general lacking, the representa¬ 
tive projectiles illustrated in Figure 2 indicate a com¬ 
plete disregard for the principles of interior and ex¬ 
terior ballistics. 42 

The sabot itself must be so designed that: 41 (1) it 
is light in weight; (2) it is strong enough to accelerate 
the projectile without breaking up in the bore, even 
when fired from a worn gun; (3) it is gas tight; (4) it 
communicates sufficient angular momentum to the 
projectile to give it full spin; (5) it is disengaged from 
the subcaliber projectile with negligible influence on 
the trajectory; (6) it holds the projectile firmly, for 
convenience in handling the ammunition; and (7) it 
can be manufactured easily and at a reasonable cost. 





Figure 2. Early sabot-projectiles. (This figure has ap¬ 
peared as Figure 1 of NDRC Report No. A-234.) 


will make it possible to fire the subcaliber projectile 
from the gun as effectively as possible. One important 
feature is stability (discussed in the next section), the 
neglect of which seems to have been responsible for 
the lack of success with the previous attempts to 
develop sabot-projectiles. 


29,2,2 Stability of Sabot-Projectiles 14 42 

The theory of the stability of subcaliber projectiles 
in general is presented in Section 8.3. As pointed out 
there, it is desirable that the stability factor be at 
least 1.5 in order to allow for variations in atmos¬ 
pheric conditions and to provide adequate damping 
of the yaw. If a gun is to be designed for shooting a 
particular projectile by means of a sabot, the twist of 
rifling can be adjusted so as to give the necessary 
stability; but if a sabot-projectile is to be shot from 
an existing gun, the stability requirements imposed 
by the twist of rifling restrict the diameter and design 
of the subcaliber projectile. 

Figure 3 shows the relationship between the dia¬ 
meter of the subcaliber projectile and the twist of 
rifling of the gun when the stability factor of the pro- 



TWIST OF RIFLING OF GUN CALIBERS PER TURN 

Figure 3. Relationship between the diameter of a 
subcaliber projectile and the twist of rifling of a gun for 
a stability factor of 1.5. The design of the projectile is 
similar to that of a standard armor-piercing one. 


CONFIDENTIAL 























































560 


SABOT-PROJECTILES 


jectile is 1.5 and its design is similar to that of a stand¬ 
ard armor-piercing projectile. This figure is based on 
the 57-mm APC projectile, M86, and would be cor¬ 
rect only if the various sizes of subcaliber projectiles 
were exactly scaled models of this particular projec¬ 
tile and if the velocities were all 2,800 fps. If another 
such graph were based on an armor-piercing projectile 
of similar design, or if account were taken of the higher 
velocities obtainable with the subcaliber projectiles, 
the differences would be relatively minor. 

If the subcaliber projectile is made of a much denser 
material, such as tungsten carbide, but is otherwise 
similar to standard armor-piercing projectiles, a 
much greater reduction of diameter is possible than 
with a steel subcaliber projectile. In practice, the 
subcaliber projectiles using tungsten carbide have 
had steel sheaths around a tungsten carbide core. 
(For example, see Figure 8). Thus it has proved prac¬ 
ticable to obtain only a part of this advantage with 
respect to stability. 

If a standard-shaped projectile of a particular di¬ 
ameter would be unstable when fired from a given 
gun by means of a sabot, it may possibly be made 
stable in this use by changing its center of gravity. 
This change can be effected by changing the shape of 
the windshield, by shortening the body of the pro¬ 
jectile, or (in the case of tungsten carbide cored pro¬ 
jectiles) by moving the core forward. However, the 
first two solutions result in greater retardation while 
the last is restricted by the geometry. 

Another possible way of increasing the stability of 
such a combination is the use of fins canted parallel 
to the rifling of the gun. This method would be espe¬ 
cially applicable in a sabot-projectile since there 
would be room in the sabot for the fins. In a few pre¬ 
liminary tests by Division 1 trouble was encountered 
in holding the fins on the projectile and so the results 
were inconclusive. 41 The Germans tried finned sabot- 
projectiles 305 ’ 387 ’ 392 during World War II, but their 
effectiveness has not yet been evaluated. 

Designs of subcaliber projectiles suitable to be fit¬ 
ted with sabots and fired from 76-mm and 90-mm guns 
with a twist of 1:32 were prepared at the Ballistic 
Research Laboratory, Aberdeen Proving Ground. 209 

29 2 3 Types of Sabot Release 

Disengagement of the sabot from the subcaliber 
projectile requires the application of a differential 
force at about the time the projectile leaves the muz¬ 
zle of the gun. Three different types of force have 


been used: the centrifugal force of the rotating sabot 
itself, the propulsion of high-pressure gas in a cavity 
at the base of the projectile, and the resistance of the 
atmosphere acting on the sabot. Two essentials are 
that the applied force not disturb the flight of the 
projectile to such an extent that its accuracy is im¬ 
paired and that the disengagement always takes place 
sufficiently close to the same point in the trajectory 
that the range of the projectile is not affected. If the 
gun is fitted with a muzzle brake, an added require¬ 
ment is that the disengagement occur in such a man¬ 
ner that there is no interference with the brake. In 
practice this meant that it be delayed until after the 
sabot-projectile had cleared the brake, for the U. S. 
Ordnance Department was unwilling to increase the 
clearance between the projectile and the baffles in the 
brake (which would decrease the efficiency), as was 
done by the British. 394 

A theoretical analysis of types of sabot release at 
the beginning of Division l’s interest in the subject 
led to the conclusion 41 that the use of centrifugal 
force to disengage the sabot was preferable. It was 
considered that it was less likely to cause perturba¬ 
tion of the trajectory because it could be applied 
symmetrically in a radial direction, whereas the other 
two forces, applied along the axis of the projectile, 
would exaggerate any yaw that happened to be pres¬ 
ent when separation occurred. Firings with 20-mm 
projectiles confirmed this hypothesis; 41 but later ex¬ 
perience with larger caliber ones showed that satis¬ 
factory accuracy could be achieved with axial release 
under certain conditions, as described below. 

Centrifugal Release 

The sabots that were released centrifugally were 
made of two principal parts: the sabot body and the 
base. It happened that the early models made both 
by the Geophysical Laboratory 41 and by the Univer¬ 
sity of New Mexico 42 had plastic bodies and metal 
bases. The plastic body, which was usually just a 
sleeve that fitted tightly over the subcaliber projec¬ 
tile, was weakened by longitudinal slots, in order that 
it would break apart after the projectile left the gun. 
In the 20-mm size (Section 29.3.1) the base was a 
solid Dural plate to which the body was threaded, 
whereas in the 75-mm sabots (Section 29.3.2) and the 
105-mm sabots (Section 29.3.5) the base was a steel 
ring screwed on the rear end of the subcaliber pro¬ 
jectile and it had radial slots so that it too was broken 
into several pieces by centrifugal force. 


CONFIDENTIAL 




DESIGN OF SABOT-PROJECTILES 


561 


When it was found desirable to substitute a light- 
alloy for plastic in the body of the 75-mm and 105-mm 
sabots (see Section 29.2.4), the body was made in 
four segments that were held together by a steel ring 
at the forward end, termed a bourrelet band. This 
ring was of such strength that it was broken by cen¬ 
trifugal force, and thus the segments of the body were 
released. 

A serious difficulty with the simple type of centrif¬ 
ugal release just described was that the sabot began 
to disintegrate as soon as it emerged from the muzzle, 
and hence would interfere with a muzzle brake. In an 
effort to prevent this action with an all-metal sabot 
for the 76-mm gun, which used a muzzle brake, sev¬ 
eral types of delayed centrifugal release were tried. 
In the simplest, the segments of the sabot body were 
held in place by a ring at the rear that was just strong 
enough to expand slowly by centrifugal action after 
it had cleared the muzzle brake, until finally it broke 
and released the segments. A preferable arrangement, 
which was adopted for the 76-mm sabot-projectile 
Model 3-76-EH described in Section 29.3.3, is de¬ 
scribed in the paragraph “Combined Axial and Cen¬ 
trifugal Release.” 

Release by Gas Pressure 

A sabot may be blown off the subcaliber projectile 
by means of compressed gas confined in a space be¬ 
tween the rear face of the projectile and the base of 
the sabot. In some experiments 100 ’ 133 this gas was 
some of the propellant gas itself, which entered the 
space through a small hole while the gas pressure was 
near its maximum value; and in other experiments 41 
it was obtained by the explosion of a small charge of 
black powder placed in the space and ignited by the 
powder gases. Neither of these methods was tried 
extensively, because the preliminary trials were un¬ 
promising, the projectiles having been very inaccurate. 
Inasmuch as the first trials of axial release sabots 
made at about the same time gave equally unsatis¬ 
factory results, whereas later efforts were successful, 
this method would seem to deserve further consider¬ 
ation. It was used successfully in combination with 
the method of axial release in a sabot designed in 
Canada for both the British 6-pounder 394 and 17- 
pounder guns. 100 

Axial Release 

In the method of release just described the sabot 


separates from the subcaliber projectile by movement 
along their common axis, but the term axial release 
has been reserved for the type of release that makes 
use of air resistance to separate them. In order to 
make this type of release practical there must be some 
provision for locking the two parts together to prevent 
separation both during handling prior to firing and 
also during passage down the bore of the gun. 

One type of lock consisted of a series of shear pins 
extending through the sabot body into the subcaliber 
projectile, so designed that they broke under set¬ 
back. 133 A more satisfactory method of locking the 
sabot to the subcaliber projectile was to have a tracer 
pocket on the latter extend through the center of the 
steel base plate of the sabot and then to attach some 
sort of retaining ring to this projection. First tried on 
a 105-mm sabot-projectile, 236 this locking means was 
later applied to other sizes and types. 

Of the different retaining rings tried, the most 
satisfactory was a split steel ring designed to open 
under the influence of centrifugal force. This type of 
retaining device was used for both the 76-mm sabot- 
projectiles of Design 3-76J (Section 29.3.3) and the 
“deep-cup” 90-mm sabot-projectile (Section 29.3.4). 
Thus, although the final separation of such a sabot 
from its subcaliber projectile is caused by differential 
air resistance, the unlocking is centrifugal. 

Earlier experiments with deep-cup sabots for 20-mm 
projectiles 41 and for 57-mm ones 42 100 were unsuccess¬ 
ful, because of large yaw after separation, whereas 
cup-type sabots developed for a 90-mm gun (Section 
29.3.4) separated successfully. The reason for this 
difference in behavior is not clear from the record. 

Combined Axial and Centrifugal Release 

In one type of 76-mm sabot-projectile intended to 
be fired through a muzzle brake, separation occurred 
in two stages. 100 First, axial drag caused the base 
plate of the sabot to move rearward with respect to 
the sabot body, which was made of segments of light 
alloy. These segments were held at their forward ends 
under a bourrelet band of steel, which was strong 
enough to resist the centrifugal thrust until the rear 
ends of the segments were released. Then the added 
thrust that occurred when the rear ends were moved 
outward by centrifugal force broke the bourrelet 
band. 

With this design also the initial locking of the sabot 
to the subcaliber projectile was very important. As is 
brought out in Section 29.3.3, 76-mm projectiles made 


CONFIDENTIAL 



562 


SABOT-PROJECTILES 


according to University of New Mexico designs per¬ 
formed satisfactorily when fired at Aberdeen Proving 
Ground, 223 whereas another group made by the Rem¬ 
ington Arms Company and also fired at Aberdeen 223 
failed, in some cases by breaking up in the bore. The 
only difference in the design of the two lots was the 
substitution of a garter spring of closely coiled piano 
wire for a split steel ring threaded to the exterior of 
the tracer pocket. The spring, which was a less secure 
type of locking arrangement, apparently permitted 
separation of the parts at the time of loading or the 
start of travel, with consequent break-up of the pro¬ 
jectile in the bore. This cause of failure was not rec¬ 
ognized at the time, 133 because of other doubtful 
factors that subsequently were better understood. 100 

29 2 4 Materials for Sabot Construction 

All the successful designs of medium-caliber pro¬ 
jectiles (See Section 29.3) have used steel bases for 
the sabots. In addition steel was used for the entire 
sabot of the cup-type developed for the 90-mm gun, 
as described in Section 29.3.4. This sabot-projectile 
has a very low ratio of weight of subcaliber projectile 
to that of the entire projectile (see Table 1, p. 567). It 
would seem, therefore, that an extensive re-evaluation 
of design factors is necessary before it can be certain 
that an all-steel sabot can be made as light as those of 
other materials. 

Light Alloys 

Both Dural (aluminum alloy) and Dowmetal (mag¬ 
nesium alloy) have been used successfully for the 
body of a sabot-projectile in the 76-mm size. 100 Dural 
was also used for the base of the sabot for 20-mm 
projectiles. 41 

These alloys are easily eroded by hot powder gases, 
and therefore it is essential that streaming of the 
powder gases over them be prevented by proper ob¬ 
turation. Furthermore, these alloys are not strong 
and hard enough to resist engraving at the bourrelet. 
To prevent it, sabots of these alloys need to have 
steel bourrelet bands, such as used with the 76-mm 
projectiles (Section 29.3.3). 100 

Plastics 

An examination of the properties of materials of 
construction led to the conclusion 54 in 1943 that or¬ 
ganic thermosetting plastics were the only nonmetallic 


materials having sufficient strength to be used for 
sabots. Sabot bodies made of this kind of plastic have 
given satisfactory perf ormance when fired soon enough 
after manufacture so that no dimensional change of 
the plastic took place. 41 - 42 After exposure to a humid 
atmosphere, however, the plastics used for 75-mm 
and 105-mm sabot-projectiles swelled to such an ex¬ 
tent that the rounds could not be fired. 100 

Tests were made at the Geophysical Laboratory 
of a number of phenol-formaldehyde plastics with 
fiber or paper filling preparatory to using them for 
some 20-mm sabots. Exposure for 3 months to ex¬ 
tremes of humidity (1% and 90%) at a temperature 
of 104 F caused linear dimensional changes of about 
+ 0.3 per cent. For a 90-mm sabot this would corre¬ 
spond to a change in diameter of + 0.01 in., which is 
larger than the tolerance permitted. The sliding 
bourrelet mentioned in Section 29.3.1 might be a 
means of allowing for such a change in diameter. 

Experimental all-plastic sabots of various designs 
were fired from both a 57-mm and a 75-mm gun. 
Most of them were machined and threaded on the 
subcaliber projectile, but a few of them were molded 
directly on it. The results were erratic, only a few 
rounds having shown satisfactory flight. 54 A critical 
examination of the work has led to the conclusion 
that “it appears improbable that these [laminated] 
plastics are strong enough to be used in all-plastic 
sabots for spin-stabilized projectiles to be fired from 
present day guns.” 100 

29 2 5 Obturation 

The experience gained during the development of 
a sabot-projectile for the 90-mm gun, M1A2 (Section 
29.3.4) confirmed the earlier experience of the British 
in the development of a sabot for the 6-pounder gun 394 
that effective obturation is very important in pro¬ 
moting the best accuracy. If parts of the sabot are 
made of Dural or Dowmetal, obturation is also essen¬ 
tial to prevent gas-washing of those parts with con¬ 
sequent failure of the projectile in the bore. 

29 3 SABOT-PROJECTILES DEVELOPED 
BY DIVISION 1 

29 3 1 Sabot-Projectiles for 20-mm Gun 

The final design of sabot-projectile developed by 
the Geophysical Laboratory 41 in tests with the 20-mm 
1 lispano-Suiza gun was of the centrifugal release type 


CONFIDENTIAL 




SABOT-PROJECTILES DEVELOPED BY DIVISION 1 


563 


(Section 29.2.3). The plastic body, which was a close 
fit on the core, was screwed to the Dural base. Type 
CW6, shown in Figure 4, had an 11-mm subcaliber 
projectile of heavy alloy (to simulate tungsten car¬ 
bide). In another type the subcaliber projectile was 
steel. A special feature of these projectiles was the 
sliding bourrelet, which was found to be essential for 
accuracy. In the early models it had been on the out¬ 
side, but the arrangement shown in Figure 4 has the 
advantage that it avoids having an exterior groove in 
which dirt might collect. Handmade projectiles of 
these designs performed reasonably well when fired 
single shot. A muzzle velocity of 4,130 fps was attained 
with the tungsten carbide core. The dispersion of the 
best rounds was about 40 per cent greater than that 
of standard ammunition fired from the same gun. 

The original objective of this program had been to 
work out a satisfactory sabot mechanism in the 20-mm 
size and then to adapt it to larger caliber guns for use 
against tanks. Accordingly, this design was adapted 
to a sabot with which to fire the 57-mm APC projec¬ 
tile, M86, from a 3-inch gun. These sabot-projectiles 
were used in firings at Aberdeen Proving Ground to 
test the armor penetration of the M86 projectile at 
high velocities, 235 ’ 242 after preliminary firings 206,232 ’ 233 
had demonstrated the feasibility of this plan. Another 
adaptation of the 20-mm design was used for the first 
design of a 105-mm sabot-projectile with a 75-mm 
subcaliber projectile. 6 Partially satisfactory results 
were obtained in preliminary firings 236 at Aberdeen 
Proving Ground. 

An attempt was made to develop a design of 20-mm 
sabot-projectile that could be manufactured by mass 
production methods, for possible use in the 20-mm 
Hispano-Suiza gun. The subcaliber projectile was the 
15-mm steel core for the caliber .60 ammunition then 
under development by Frankford Arsenal. For the 
sake of simplifying the manufacture of the plastic 
sabot body no sliding bourrelet was used. It was not 
possible to mold the plastic body with a small enough 
tolerance, and difficulty was experienced in fitting the 
Dural bases to the sabot bodies. Nevertheless a small 
lot of these projectiles (Model C5b) were prepared 
for preliminary firing tests at Aberdeen Proving 
Ground to determine the exterior ballistics 207 ’ 237 of 
these projectiles and the interior ballistics 234 of the 


e These projectiles were supplied by the Geophysical Lab¬ 
oratory. Because better facilities for an extensive development 
existed at the University of New Mexico, the 105-mm project 
was continued there, with the results described in Section 
29.3.5. 


gun firing them. It was concluded that the poor effi¬ 
ciency in armor penetration that was observed was 
due to slippage between the'subcaliber projectile and 
the base. Furthermore, as mentioned in Section 29.2.4, 
considerable variation in dimensions was observed 
when various plastics (intended for the sabot body) 
were exposed to different conditions of humidity. 100 
Thereupon this project was abandoned. 

MACHINABLE ALLOY 



Figure 4. 20-mm sabot-projectile, Model CW6, with 

sliding plastic bourrelet and dense core, employing cen¬ 
trifugal release. (This figure has appeared as part of Fig¬ 
ure 9 in NDRC Report No. A-233.) 

29 3 2 Sabot-Projectiles for 

75-mm Gun and Howitzer 

The sabot-projectiles for the 75-mm tank gun, M3, 
also were of the centrifugal-release type (Section 
29.2.3) with a plastic body that was a press-fit on the 
subcaliber projectile. The base of the sabot was a 
steel ring screwed to the base of the subcaliber pro¬ 
jectile. These projectiles Avere developed by the Uni¬ 
versity of New Mexico 42 ’ 100 following preliminary 
experiments with sabot-projectiles fired from a Navy 
6-pounder gun (caliber: 57 mm). 

The core of Model 28-75D, shown in Figure 5, was 
a slightly modified 57-mm APC projectile, M86, 
weighing 7 lb. A lot of these projectiles was fired at 
Aberdeen Proving Ground 238 at a muzzle velocity of 
about 2,800 fps compared with 2,030 fps for the stan¬ 
dard 75-mm APC projectile, M61. Although the ac¬ 
curacy and other features of performance were satis¬ 
factory, the using Services decided that at that time 
(just after the Normandy invasion) there was no 
tactical use for this projectile for the 75-mm gun, 
especially in view of the expectation that tanks would 
soon be equipped with 76-mm guns having a muzzle 
velocity of about 2,600 fps for a 15-lb projectile. 100 


CONFIDENTIAL 

















564 


SABOT-PROJECTILES 


About 6 months later a slight modification of this 
design was prepared for use in the 75-mm pack how¬ 
itzers, M1A1, M2, and M3. Preliminary firings were 
made at Aberdeen Proving Ground (July 1944) with 
satisfactory results. A number of these projectiles 
were forwarded to the Armored Board, Fort Knox, 
Kentucky, for test in comparison with standard am¬ 
munition. By that time it had been learned that 
plastic sabots for the 105-mm howitzer (Section 29.3.5) 
had swelled from exposure to a humid atmosphere, 
and this 75-mm ammunition was examined. It also had 



Figure 5. 75-mm sabot-projectile, Model 28-75D, 

with plastic sabot body and steel sabot base ring that 
separate centrifugally. (This figure appeared in a prog¬ 
ress report from the University of New Mexico on 
Contract OEMsr-668.) 

swelled to such an extent that it could not be loaded 
in the gun. Drawings were made of a modified design 
in which light alloy would be substituted for the 
plastic, 100 but these projectiles were not constructed. 

Previously an entirely different type of all-metal 
75-mm sabot-projectile using the 57-mm projectile, 
M86, for the subcaliber part had been made for Div¬ 
ision 1 by the Budd Wheel Company and three rounds 
had been fired at Aberdeen Proving Ground. 239 The 
test was inconclusive and no further development 
was made. 

29 3 3 Sabot-Projectiles for 76-mm Gun 

When the Army Ordnance Department requested 
Division 1 to develop a sabot-projectile with tungsten 
carbide core for the 76-mm gun, M1A2, it was de¬ 
cided to make this sabot an all-metal one. The design 


was complicated by the fact that the projectile had to 
pass through a muzzle brake with only a slight clear¬ 
ance. 

In order to satisfy the latter requirement, the de¬ 
sign finally developed at the University of New Mex¬ 
ico, 100 which is shown in Figure 6, employed the 
delayed centrifugal release described at the end of 
Section 29.2.3. The core was a tungsten carbide one 
similar to that used by the British in a sabot-projectile 
for the 17-pounder gun (caliber: 3 in.). Determina¬ 
tions of velocity vs time of flight and of range vs time 
of flight were made at the New Mexico proving- 
ground by the method of tracer photography 57 de¬ 
scribed in Section 8.6. 

A first lot of these projectiles (Design 3-76EH) was 
fired at Aberdeen Proving Ground with satisfactory 
results 223 at muzzle velocities slightly above 3,600 fps. 
X-ray photographs 210 were made of some rounds to 
study the separation of the sabot from the subcaliber 
projectile. The design was modified in a number of 
minor details, the most important of which was the 
substitution of a steel retaining ring (Section 29.2.3) 
for a Dowmetal one. Projectiles of this design (3-76J) 
were also fired at Aberdeen Proving Ground with 
satisfactory results. 223 

Subsequently the Remington Arms Company mod¬ 
ified the design in an attempt to develop a model 
suitable for mass production. 100 ’ 133 As mentioned in 
Section 29.2.3, this model when fired at Aberdeen 
Proving Ground 223 (where it was designated “76-mm, 
T23”) failed in the bore, presumably because the 
garter spring that had been substituted for the steel 
retaining ring did not hold the parts together at the 
beginning of travel. Although a number of successive 
modifications were fired at Aberdeen Proving 
Ground, 223 in none of them was this fault corrected, 
and all were erratic in their behavior. 

An entirely different design (76-48-R2) was tried 
later. 543 Although it was patterned after the later 
designs of 90-mm deep-cup sabots (Section 29.3.4), it 
was inaccurate. The reason was not determined. 

29 3 4 Sabot-Projectiles for 90-mm Gun 

The experience gained with the 76-mm sabot- 
projectile (Section 29.3.3) was applied to the devel¬ 
opment of a sabot-projectile for the 90-mm gun, 
M1A2, which also used a muzzle brake. The first 
such projectile, constructed by the Remington Arms 
Company 133 from a design prepared by the Univer¬ 
sity of New Mexico, 100 was similar to the latter’s 


CONFIDENTIAL 



SABOT-PROJECTILES DEVELOPED BY DIVISION 1 


565 


76-mm projectile described in the previous section, 
in that it had a segmented sabot body of Dowmetal. 
To it was fitted a copper obturator. This model failed 
when fired at Aberdeen Proving Ground, 222 as did a 
number of successive modifications of it, 222 which 
were designated as the “90-mm, T32 series.” The 
reason for these failures was later discovered to be 
the separation of the sabot from the subcaliber 
projectile in the bore at the beginning of travel, as 
had happened also with the Remington 76-mm 
models. 100133 

An entirely different type of sabot (the “deep-cup” 
type described in Section 29.2.3) was then developed 
for this gun by the Remington Arms Co. 133 At first 
the firings were carried out at Aberdeen Proving- 
Ground, 222 where these projectiles were designated 
as the “90-mm, T38 series,” and later at a private 
range at Pine Camp, N. Y. 

The final design (90-56-R13), shown in Figures 7 
and 8, used a subcaliber projectile that consisted of a 
6-lb tungsten carbide core contained in a steel sheath. 
Reference should also be made to Figure 1, which is 
a schematic representation of this design. On the 
basis of a single shoot of 5 rounds from a new gun, 
this was by far the most accurate sabot-projectile of 
any size developed by Division 1. The five shots, 
which had a muzzle velocity of 3,383 fps, struck with¬ 
in a rectangle 1.3 ft wide and 0.8 ft high at 1,000 yd. 
This accuracy is essentially the same as that obtained 
with standard 90-mm ammunition. A corresponding 
model (90-58-R84) was made up with an 8-lb core, 


but there was time to fire only two such projectiles 
before the close of the project. Their performance 
was satisfactory. 133 

The high degree of accuracy achieved with these 
90-mm projectiles was attributed to several features 
of the design: 

1. The sabot was long enough relative to its diam¬ 
eter to avoid any tendency toward cocking in the 
bore. 

2. An effective rubber obturator prevented escape 
of powder gases past the sabot at the beginning of 
travel, which might have heated and thereby weak¬ 
ened the bourrelet or which might have pushed the 
projectile out of line with the axis of the bore. 

3. A pair of sintered iron rotating bands had ade¬ 
quate strength to spin the projectile. (It should be 
noted, however, that there was no opportunity to test 
this final model in a worn gun tube.) 

4. The subcaliber projectile was firmly attached 
to the sabot, which prevented lateral movement dur¬ 
ing loading and during the beginning of travel and 
which also insured that full spin was imparted to the 
subcaliber projectile. 

5. The middle bearing (which was just to the rear 
of the center of gravity) in conjunction with the 
bourrelet ring maintained concentricity of the sub¬ 
caliber projectile with respect to the sabot and hence 
with respect to the axis of the gun. 

6. When separation did occur, considerable clear¬ 
ance resulted from only Yi in* of relative motion 
along the axis; and hence there w r as no appreciable 



Figure 6. 76-mm sabot-projectile, Design 3-76EH, with light-alloy sabot body and steel base plate that separate by 

axial and delayed centrifugal release. (This figure has appeared in NDRC Report No. A-461.) 


CONFIDENTIAL 






566 


SABOT-PROJECTILES 


interaction between the sabot and the subcaliber 
projectile. 

During the design and construction of the succes¬ 
sive models of 90-mm sabot-projectiles attention was 
paid to production engineering. As a result, at the end 
of the project it was possible to make recommenda¬ 
tions for the equipment and manpower required to 
produce these projectiles at a rate of 2,000 a day. 134 



Figure 7. 90-mm all-steel sabot-projectile, Model 

90-56-R13, with tungsten carbide core, employing axial 
release. (This photograph has appeared in NDRC Re¬ 
port No. A-461.) 

29,3,5 Sabot-Projectiles for 105-mm 
Howitzer 

The 105-mm howitzer, M3, is a lightweight weapon 
developed for the use of airborne troops. Ordinarily 
it fires a 33-lb HE shell at a muzzle velocity of 1,020 
fps or less, depending on the range desired. The use of 
a sabot-projectile seemed to offer an opportunity to 
provide a moderately high-velocity AP shot for the 


same gun and also a smaller HE shell having a high 
enough muzzle velocity to be usable against “point” 
targets. 

The two sabot-projectiles developed at the Univer¬ 
sity of New Mexico for this howitzer were of the same 
general design and differed chiefly in the subcaliber 
projectiles. 1 One was the 57-mm APC projectile, 
M86; the other was the 3-in. HE shell, M42A1. For 
the former it was necessary to determine the stability 
factor, using the experimental method currently in 
use by the Ballistic Research Laboratory, Aberdeen 
Proving Ground, 177 and then to shorten the wind¬ 
shield on the projectile in order to increase the stabil¬ 
ity factor to a value of 1.5. 100 

The sabot used for these projectiles was similar to 
that of the 75-mm Model 28-75D (Figure 5). Both 
the AP projectiles (Model 2-105R) and the HE pro¬ 
jectiles (Model 3-105B) were fired at Aberdeen Prov¬ 
ing Ground with satisfactory results. 224 Then about 
100 of each type were sent to the Infantry Board, 
Fort Benning, Georgia, for Service acceptance tests. 
There difficulty was encountered in loading the pro¬ 
jectiles in the gun because of swelling of the plastic 
sabot bodies. After the excess material had been 
removed by machining, some of the projectiles were 
fired. The results were satisfactory, except that the 
gun hop was excessive. Some AP rounds were also 
fired from the heavier howitzer, M2. On the basis of 
all these tests it was decided that there w^as no further 
interest in the 105-mm sabot-projectile that used the 
3-in. HE shell for its subcaliber projectile. There was 
some interest in the substitution of light alloy for 
plastic in the 57-105-mm combination, but the devel¬ 
opment of sabot-projectiles for the 76- and 90-mm 
guns pre-empted the available facilities. 

29 4 CONCLUSION 

The investigation of sabot-projectiles by Division 
1 has involved experiments with a wide variety of 
designs. Whereas this field previously had been re¬ 
garded by ordnance specialists as unprofitable, these 
experiments have demonstrated the great potential 
usefulness of sabot-projectiles. Although none of 
them reached the final stage of test under combat 
conditions, there has been accumulated a wealth of 
basic design data. The outstanding conclusions of the 
work have been summarized in this chapter; in ad- 


f A different type of sabot-projectile for the 105-mm howitz¬ 
er is described in Section 29.3.1. 


CONFIDENTIAL 








CONCLUSION 


567 



Figure 8. 90-mm all-steel sabot-projectile, Model 90-56-R13, with tungsten carbide core, employing axial release. 

(This drawing has appeared in NDRC Report No. A-461.) 


dition to which there are many useful details in the 
original reports cited. 

The characteristics of the final designs of sabot- 
projectiles of different sizes, which have been dis¬ 
cussed in Section 29.3, are given in Table 1. The best 
accuracy was achieved with the 90-mm one. None of 
these projectiles was constructed and fired in suffici¬ 
ently large quantities to determine how serious would 
be the variation in performance caused by variations 


in the dimensions of the projectiles from large-scale 
production. One of the chief purposes of the pilot- 
plant operations planned for the 90-mm sabot-pro¬ 
jectiles was to make available a quantity of projectiles 
for such tests. This phase of the work was not carried 
out because of the demobilization of OSRD. 

Each of the projectile designs listed in Table 1 has 
one or more features that is worth further study. Thus 
the delayed centrifugal release mechanism of the 


Table 1. Characteristics of final designs of sabot-projectiles developed by Division 1 contractors. 


Caliber (mm) 
Subcaliber diam. (mm) 
Core diameter (mm) 
Core material* 
Projectile model 
Gun model 
Type of release f 
Weight of subcal. (lb) 
Total weight (lb) 
Weight ratio 
Muzzle velocity (fps) 
MV stand, proj. (fps) 
Wt. stand, proj. (lb) 
Chamber pr.§ (psi) 
Twist of rifling 
Number rounds fired 
Bibliographic ref. 


20 

75 

15 

57 

15 

57 (M86) 

S 

S 

C5b 

28-75D 

M1A2 

M3f 

C 

C 

0.147 

7.00 

0.207 

8.38 

0.71 

0.84 

3,310 

2,800 

2,615 

2,030 

0.36 

14.96 

48,000 

36,000 

1:25.58 

1:25.58 

ca. 50 

69 

207 

41,100 


76 

90 

55.9 

55.9 

38.1 

45.0 

WC 

WC 

3-76EH 

90-56-R] 

M1A2 

M1A2 

A-C 

A 

5.77 

8.06 

7.97 

12.50 

0.72 

0.64 

3,650 

3,880 

2,600 

2,650 

15.44 

24.11 

43,000 

38,000 

1:32 

1:32 

8 

5 

100, 223 

133 


105 

105 

57 

76 

57 (M86) 

76 (M42A1) 

S 

S 

2-105R 

3-105B 

How, M3 

How, M3 

C 

C 

7.207 

12.7 

11.1 

15.6 

0.66 

0.81 

2,660 

1,980 

1,020 

1,020 

33.00 

33.00 

30,000 

30,000 

1:20 

1:20 

10 

10 

100, 224 

100, 224 


* S = steel; WC = tungsten carbide. 

t This projectile was also fired from the 75-mm pack howitzer. 

t C = centrifugal release; A = axial release; A-C = axial and delayed centrifugal release. 
§ Rated maximum pressure for the gun. 


CONFIDENTIAL 
































































568 


SABOT-PROJECTILES 


76-mm projectile, if used in conjunction with a good 
obturator (such as w T as used later in the 90-mm pro¬ 
jectile) and adequate means to hold the parts to¬ 
gether, might be very effective for projectiles that 
need to pass through a muzzle brake. For projectiles 
that do not need to pass through a muzzle brake, the 
simple Design 28-75D with light alloy substituted for 
plastic in the sabot body should be satisfactory, es¬ 
pecially if an obturator is added. 

If at some time in the future a plastic is developed 
that has greater dimensional stability than those 


tried, it would be worth while to reopen the question 
of the use of plastics in sabots. Perhaps the sliding 
bourrelet of the 20-mm projectile added to Design 
28-75D would increase the accuracy of this type. 

Because the energy imparted to the discarded parts 
of a sabot-projectile is wasted, it is desirable to keep 
the weight of the sabot to a minimum. By making 
use of the total experience with sabot-projectiles that 
is now available, it should be possible to develop 
designs by which this can be accomplished without 
sacrificing reliability of performance. 


CONFIDENTIAL 



Chapter 30 

TAPERED-BORE GUNS AND SKIRTED PROJECTILES* b 

By Edwin L. Rose c 


301 INTRODUCTION 

T he tapered-bore gun, firing a skirted projectile, 
offers one of several means for the attainment of 
hypervelocity without exceeding currently existing 
limits on powder-gas pressures and temperatures. It 
produces this hypervelocity by driving a projectile of 
reduced mass with the same force as that applied by 
the powder gas to the base of a standard weight pro¬ 
jectile. In common with the standard-bore gun firing 
a sabot-projectile (Chapter 29) this combination offers 
a marked advantage over the standard gun firing a 
lightweight projectile of equal weight but of stand¬ 
ard caliber, since the retardation in flight of the spe¬ 
cial projectiles is less by a factor which is roughly the 
square of the ratio of the diameters. The combination 
of tapered-bore gun and skirted projectile possesses 
an appreciable advantage over the sabot arrange¬ 
ment in that all the parts are attached to the projec¬ 
tile so that the entire muzzle momentum is available 
for overcoming air resistance. The importance of 
these advantages is shown strikingly in Section 33.2 
where the relative performances of different types of 
hypervelocity guns, in various types of service, is 
analyzed in some detail. 

In the light of the recently developed facts, it 
appears that the use of tapered-bore guns is not merely 
an expedient for the attainment of hypervelocity 
with standard pressures, but that regardless of future 
improvements in the velocity of projectiles of stand¬ 
ard caliber important gains will always be attainable 

a The term “skirted” is used here in lieu of “deformable” to 
avoid confusion with the prior use of the term in connection 
with terminal ballistics to designate soft or plastic projectiles 
as distinguished from rigid, nonplastic, or hard projectiles. 

b The tapered-bore gun described in this chapter was devel¬ 
oped for Division 1, NDRC, by the Jones and Lamson Machine 
Company under Contract OEMsr-467. The development of the 
skirted projectiles for this gun was first undertaken by the 
Bryant Chucking Grinder Company under Contract OEMsr- 
534, and was later continued by the Jones and Lamson Ma¬ 
chine Company. This chapter is based on the combined final 
report 128 under those two contracts. The history of the develop¬ 
ment of tapered-bore guns had been covered in an earlier 
report. 2 

c Member, Division 1. (Present address: Consulting Engi¬ 
neer, Jones and Lamson Machine Company, and Bryant 
Chucking Grinder Company, Springfield, Vt.) 


from the addition of the tapered bore feature. This ob¬ 
servation is in direct opposition to military consensus. 
It therefore requires critical consideration in order to 
check its validity and to make sure that an oppor¬ 
tunity for a substantial improvement in the effective¬ 
ness of weapons is not overlooked until too late. 

The development of a 57/40-mm tapered-bore gun 
and skirted projectile was originally undertaken by 
Division 1, NDRC, for the purpose of making avail¬ 
able to our Armed Forces designs, manufacturing 
equipment, and methods d for producing tapered-bore 
guns and ammunition of the Gehrlich type. 6 Soon 
after the work was undertaken it appeared that the 
production of such guns was unnecessarily tedious 
and expensive in the consumption of man-hours of 
critical labor. It was thought that if suitable projec¬ 
tiles designed to withstand the shock of extremely 
rapid deformation could be devised, the Janacek type 
of gun, which uses a standard tube with a short ta¬ 
pered adaptor attached to the muzzle, would be much 
more practical. In addition to simplifying the prob¬ 
lem of manufacture its use would open the way to 
conversion of existing guns for the firing of skirted 
projectiles. 

Tests on a number of tubes both rifled and unrifled, 
with tapers of different lengths, showed that it was 
practicable to drive a suitably designed projectile 
through an 8-in. long taper and thereby reduce the 
diameter of the fins from 57 mm to 40 mm at a veloc¬ 
ity in excess of 5,000 fps. It was further found that 
with this short taper the rifling could be limited to 
the standard bore section without appreciable loss of 

d A suggestion 528 of a method for rifling tapered bores, con¬ 
sidered in the early studies of this project, came from V. 
Bush. 

e Earlier, the Division had sponsored some work at the 
National Bureau of Standards on a method of static testing 
skirted projectiles. 4 Those tests had led to the design of a 
series of 37/28-mm skirted projectiles of a so-called hour-glass 
shape, which were fired at Aberdeen Proving Ground from a 
37-mm gun, M3, to which a Littlejohn adaptor of the British 
design 425 had been added. Those firings were made by the 
Army Ordnance Department as part of its own program 218 on 
the development of hypervelocity projectiles for firing from 
existing guns. The results 217 ' 218 ’ 219 ’ 227 ’ 228 * 229 ’ 230 were communi¬ 
cated to the Jones and Lamson and Bryant companies and 
were useful in the early planning of the 57/40-mm project. 
(Editor’s note.) 


CONFIDENTIAL 


569 




570 


TAPERED-BORE GUNS AND SKIRTED PROJECTILES 


spin of the projectile before emergence. All doubts as 
to the preferability of the Janecek adaptor were thus 
eliminated and this design was adopted. 

In order to carry through a completed design, the 
work was concentrated on the development of a 
tungsten carbide-cored projectile suitable for anti¬ 
tank use. The high-explosive projectile was entirely 
neglected. Nevertheless some of the greatest poten¬ 
tialities of the tapered-bore gun and skirted projectile 
lie in their use against rapidly moving, maneuvering 
targets where higher velocity greatly increases the 
probability of an effective hit. The problem of pro¬ 
ducing a satisfactory skirted projectile has been 
solved in a manner applicable to either armor-pierc¬ 
ing or high-explosive use. The main problems remain¬ 
ing to be solved for skirted high-explosive projectiles 
are the achievement of suitable stability in flight (to 
insure accurate flight characteristics), and the attain¬ 
ment of good fragmentation characteristics. These 
problems are hardly more difficult than those en¬ 
countered in the case of the armor-piercing projectile 
and they do not appear to present any insurmount¬ 
able obstacles. 

The work done on the 57/40-mm development 
appears to provide a satisfactory basis for the design 
of tapered adaptors and skirted projectiles suited to 
any velocities readily attainable with guns and pow¬ 


ders available within the visible future. Projectiles of 
the type developed have been successfully fired at 
velocities as high as 5,200 fps. It seems quite reason¬ 
able to expect that the methods outlined herein may 
produce materiel capable of successful firing at vel¬ 
ocities approaching 10,000 fps. The attainment of 
this end must, however, await the full application of 
what we have learned regarding erosion control. It is 
further dependent upon additional improvements in 
structural design to take advantage of improved 
erosion control. 

302 57/40-MM TAPERED-BORE GUN AND 

SKIRTED PROJECTILE 

As finally developed, the 57/40-mm tapered-bore 
gun comprises a tapered muzzle adaptor attached by 
a screw thread to the muzzle end of a Canadian Mk 
III 57-mm gun. The overall length of the adaptor is 
34 in.; the tapered section is 8% in.; the straight 
muzzle section, of 40-mm diameter, is 21 in.; and the 
remainder is taken up by the attachment section. 
Figure 1 is a photograph of a complete 57-mm Mklll 
mount with the tapered muzzle adaptor in place on 
the gun. This figure also shows two pieces of assem¬ 
bled ammunition at the side of the gun. The details 
of the adaptor are shown in Figure 2. The principal 



Figure 1. 57/40-mm hypervelocity tapered-bore gun, Jones and Lamson final design, on standard 57-mm field carriage. 

(This figure has appeared as Figure 54 in NDRC Report No. A-456.) 


CONFIDENTIAL 







57/40-MM TAPERED-BORE GUN AND SKIRTED PROJECTILE 


571 



(8) ^-DRILLED HOLES AS SHOWN 


MUZZLE ADAPTORS N0-9AND12 



TYPICAL DRILL LAYOUT ALL OTHER DIMENSIONS SAME AS NO. 13 (BELOW) 


SECTION AO B 


3 i 
0 16 


-16 ||- STRAIGHT 0 D 


T 

1 


3 _ 

8 



MUZZLE ADAPTOR NO.13 

Figure 2. Muzzle adaptors for 57/40-mm tapered-bore gun. (This figure has appeared as Figure 12 in NDRC Report 
No. A-456.) 


features of this design are the means of attachment, 
the tapered transition from initial to emergent cali¬ 
ber, the straight minor-caliber muzzle section, and 
the vent holes at the base of the taper. The vent holes 
serve to suppress the powder-gas shock wave which 
tends to form when the projectile enters the taper, 
and which, if unsuppressed, causes destruction of the 
cartridge case. 

The taper acts as a squeezing die, reducing the 
diameter of the projectile to its emergent size as it 
travels through this section. The straight section 
serves to set the final dimension of the projectile, to 
center its geometrical and spin axes, and to keep down 
the yaw at the instant the projectile leaves the gun. 

The attachment section is provided with a thread 
for securing the adaptor to the gun and with bearing 
surfaces to center its axis with the axis of the gun. 
There are also a seal surface and vent holes to keep 
the powder pressure from developing its full value in 
the clearance space between the end face of the gun 
tube and the base of the adaptor, thereby reducing 
the necessary strength of the securing thread. The 
securing thread is a left-hand thread. It is necessary 
only to screw the adaptor on by hand since the spin 
friction tightens it. Surprisingly, the adaptor can 


generally be removed by hand after firing numerous 
shots. 

The 57/40-mm skirted projectile is a tungsten 
carbide-cored, armor-piercing projectile, consisting of 
six parts. It is provided with a front flange or skirt to 
center the front end in the bore, and with a rear skirt 
to serve as an obturator, sealing the bore and trans¬ 
mitting the powder force to the projectile, and cen¬ 
tering the base of the projectile in the bore. 

Figure 3 is a photograph of a cut-open assembled 
projectile, while Figure 4 is an exploded view of the 
projectile with the various parts identified. The nose 
point is an aluminum or magnesium part providing a 
sharp point for minimizing the shock air wave resis¬ 
tance, and a body of soft material to minimize the 
impulse tending to deflect the projectile at the in¬ 
stant of impact against inclined armor. 

The windshield has a smooth ogival exterior con¬ 
tour to keep down air resistance. It is made of steel 
and together with the carrier helps to support the 
core against fracture on impact against inclined armor. 

The nose pad is of aluminum. It is machined with 
projecting corners where it bears on the nose of the 
core to permit some plastic flow when the windshield 
is screwed tightly against the core. The nose pad 


CONFIDENTIAL 











































































572 


TAPERED-BORE GUNS AND SKIRTED PROJECTILES 



Figure 3. 57/40-mm skirted projectile, Jones and 

Lamson final design: Quarter sectioned; two-thirds 
actual size. (This figure has appeared as Figure 62 in 
NDRC Report No. A-456.) 

serves the dual purpose of insuring a tight assembly, 
in which the core does not shift as it experiences the 
accelerations which occur in the gun, and of “soften¬ 
ing’’ the impact when the core strikes armor. The 
actual nature of this latter function is not understood, 
but it is an observed fact that armor-piercing cores 
show less tendency to shatter when fitted with such 


a pad. The container, and the skirts which form an 
integral part of it, are critical features of design. Suc¬ 
cess of a skirted projectile design depends on the 
choice of suitable material and suitable contours and 
proportions of its parts. The skirts are deformed from 
their initial to their final shapes in a time of the order 
of one to two ten-thousandths of a second. At the 
rates of strain which this deformation involves, al¬ 
most all materials show some degree of brittleness. 
The strain work at fracture is greatly reduced. Duc¬ 
tile materials must be used, sharp corners must be 
avoided, and parts subjected to bending must be 
curved and proportioned so that bending strain is 
distributed over substantial linear intervals. 

The design of the rear skirts presents a difficult 
compromise between providing enough material to 
support the powder pressure and keeping thickness of 
material down so as to minimize deformation forces 
and friction. Figure 5 shows a fully deformed, skirted 
projectile in flight. 

Since small guns, especially those in antitank serv¬ 
ice, are seldom provided with director control, the 
use of tracer ammunition is generally required. The 
tracer pocket devised for the skirted projectile is 
simple, but seems to possess advantages and novel 
features which warrant its adoption for other ord¬ 
nance use. 

30 3 DESIGN THEORY FOR TAPERED 
GUNS AND SKIRTED PROJECTILES 

30 3 1 Design of Skirts 

Early deformable projectiles were provided with 
conical skirts machined integrally with the body of 
the projectile. The line of junction between the outer 
surfaces of the skirt and projectile was a point of 
strain concentration during the final stages of the de¬ 
formation process. At high rates of strain even the 
ductile, low-carbon steels show a large reduction of 
elongation and a consequent trend toward brittleness. 
Beyond a certain critical muzzle velocity the skirts 
showed a tendency to separate from the body at this 



TRACER CARRIER OR CONTAINER 

POCKET 



CORE 


NOSE WINDSHIELD NOSE 

PAD POINT 


Figure 4. Exploded view of skirted projectile containing a tungsten carbide core to be fired from a tapered-bore gun. 


CONFIDENTIAL 








DESIGN FOR TAPERED GUNS AND SKIRTED PROJECTILES 


573 


region of high strain. This is undoubtedly a mani¬ 
festation of the reduced elongation. 

To avoid separation, skirts should be designed with 
a curved cross section. The tangent to the neutral axis 
of the section, at the point of juncture with the body, 
should be parallel to the axis of the projectile. If this 
design is followed, and if the arc of curvature is of 
suitable radius, the bending strains developed during 
deformation is distributed and conditions conducive 
to fracture are avoided. 

In addition to centering the projectile in the bore, 
the rear skirt serves as an obturator and must support 
the powder pressure. The curved section which serves 
to distribute strain also helps to reduce the skirt thick¬ 
ness necessary to support the pressure, by subjecting 
the material to nearly pure compression and eliminat¬ 
ing bending stress. In this regard the ideal section is 
one in which the curvature increases from the base of 
the projectile to the point of bearing on the bore. 
However, the gains to be made by this further re¬ 
finement do not compensate for the increased diffi¬ 
culty of tooling. 

The skirt sections visible in Figure 2 are the result 
of extensive trial and error design, guided by the con¬ 
siderations outlined above. 

30 3 2 Stresses in Skirts 

In addition to the stresses developed by the pow¬ 
der pressure and by the deformation process, the 
skirts are subject to other tensile stresses induced by 
deformation forces, by friction, and by shock as the 


projectile enters the taper. The deformation is for the 
most part by plastic compression and does not pro¬ 
duce tension. However, since the skirts are pulled 
through the taper by the body of the projectile, the 
work of deformation of the skirt material and the 
work in overcoming friction are developed by a ten¬ 
sile force. Furthermore, the skirts experience change 
of velocity on entering and leaving the tapered bore. 
There is first a loss of momentum on impact and a 
recovery when leaving the taper. The recovered 
momentum is supplied by a tensile force. This force 
appears to be the controlling one in determining- 
design. 

The bending stresses resulting from the deforma¬ 
tion process have been satisfactorily controlled by 
the curved skirt design outlined in Section 30.3.1. 
The remaining tensile stresses, developed by the work 
of plastic deformation, by friction, and by shock, ap¬ 
pear to be determined almost entirely by the design 
of the gun tube. To the extent that they can be modi¬ 
fied by projectile design, the effects are minor. 

Tensile stresses in skirts induced by the work of 
deformation and by friction have been subjected to 
an approximate analysis and the shock stresses have 
been estimated. The analysis is based on the follow¬ 
ing assumptions: 

1. Impact between skirt and wall of taper is in¬ 
elastic. 

2. Portions of skirt which have not engaged wall 
retain their original contour. 

3. Portions of skirt which have engaged wall con¬ 
form to wall. 



Figure 5. Projectile No. 451 after deformation: Taken in flight, using microflash unit; velocity 4,650 fps. (This figure 
has appeared as Figure 20 in NDRC Report No. A-456.) 


CONFIDENTIAL 




574 


TAPERED-BORE GUNS AND SKIRTED PROJECTILES 


4. No portion of skirt suffers compressive or tensile 
deformation parallel to the neutral axis of its section. 

On the basis of these assumptions, higher stress 
values were obtained than would result from any 
actual process of deformation. 

If the skirt is given any initial contour in an axial 
plane, subject to the limitation that at all points the 
tangent to the neutral axis of the section of the skirt 
makes a greater angle with the projectile axis than 
does the tangent to the tapered bore, then, when the 
skirt engages the taper, it moves over the surface of 
the taper with a sliding velocity, V s . This velocity is 
given by equation (1), 

V 8 = 7 0 (cos P — sin p tan a), (1) 

in which V 0 is the velocity of the projectile, a is the 
angle of intersection of the skirt surface with the bore 
surface, and p is the angle between the axis and the 
tangent to the taper at the same point. 

In the case where a and p are constant, V s is con¬ 
stant from the start of engagement until the base of 
the skirt enters the throat of the taper. V 8 is always 
less than V Q if a is positive. Consequently, when the 
skirt enters the throat its velocity must be abruptly 
increased from V a to Vo and a wave of stress supply¬ 
ing the necessary change in momentum of the skirt 
starts at the base and moves toward the rear of the 
skirt. The propagation of this wave is complicated by 
the nonuniform transverse section of the skirts and 
by surface friction. However, a lower limit to its mag¬ 
nitude can be calculated by assuming a uniform sec¬ 
tion and neglecting friction. The change of velocity 
AV occurring at the throat is given by equation (2). 

AV = Fo — V a = F 0 (l — cos (3 + sin /3 tan a). (2) 

It can readily be shown that this sudden change of 
velocity is accompanied by a tensile stress wave of 
amplitude T 8f expressed by equation (3), 

r. = A Vylf 

V ~Eb 

-^-(1 — cos P + sin p tan a), (3) 

in which E is Young’s modulus for the material, 8 its 
density and g the acceleration of gravity. If To is the 
tensile strength of the material, the allowable change 
of velocity AV 0 is given by equation (4), 

AVo < r »\/S’ (4) 

and the critical limits on taper and skirt angles are 


given by equation (5), 

1 - cos P + sin p tan a < (5) 

which for small values of p and a reduces to the ap¬ 
proximate form of equation (6). 

. / sin p . \ To I g , . 

In tests with 57/40-mm skirted projectiles, fired 
through tapered tubes of different lengths of taper, it 
was found that, at a muzzle velocity of 4,500 fps, pro¬ 
jectiles could be fired successfully through an adaptor 
which tapered from 57 mm to 40 mm in 8 in. of length, 
but that the skirts separated in an adaptor with a 
4-in. taper. The half angles of these two cones are 
respectively 0.04 radians and 0.08 radians. For the 
above velocity, and a skirt material having To = 
100,000, 8 = 0.28, E — 30,000,000, the left-hand side 
of expression (6) becomes <0.13. 

. / sin p . \ 

sm p l -—g-r sm a J< 

100,000 / 32.2 X 12 ~ 

4,500 X 12 \3 X 10 7 X 0.28 

The value of a is not known for these tests. However 
it was small. For a. — 0 the limiting value of P is 0.16 
and for a = p the limiting value is 0.09. When it is 
remembered that friction and the work of plastic de¬ 
formation have been omitted in this computation the 
discrepancy is seen to be in the right direction. It is 
possible that a was more of the order of 2 p, in which 
case the limiting value of P would be 0.07 radian. 
This value is between the 0.04-radian and 0.08-radian 
values indicated above. 

The tensile stress at the base of the skirt, result¬ 
ing from friction and plastic working as the skirt 
enters the throat, is given by equation (7) 

rri _ PtL (sin p + <j> cos P) m 

fw r(l — 0 sin p) K ) 

in which T/ w is the tensile stress in question, P is the 
compressive stress for plastic working, r is the radius 
of the throat, t is the thickness of the skirt, L the 
length of skirt, and <j> the coefficient of friction. This 
stress at the instant of shock is determined from 
equation (3), and the total stress is given by the sum 
of equations (3) and (7). For the materials used in the 
present designs, P is in the neighborhood of 100,000 
psi, t is about 0.125 in. for the front skirt and 0.250 in. 
for the rear skirt, and L is about 0.7 in. for the front 


CONFIDENTIAL 







DESIGN FOR TAPERED GUNS AND SKIRTED PROJECTILES 


575 


skirt, and 1.14 in. for the rear skirt. For P = 0.08 
radian and with a coefficient of friction of 0.05, 
Tf W = approx. 1,500 psi for front skirt; T fw = approx. 
6,000 psi for rear skirt. These values are of the order 
1.5 per cent to 6 per cent of the shock stress com¬ 
puted above, and can therefore have little influence 
on the critical angle. Apparently the shock stress is 
all important in determining the limiting velocity of 
a tapered-bore gun. 

Since a is limited to positive values, it is apparent 
that little can be done to the design of the skirt to 
change the velocity limit. The only effective control 
is by the adjustment of p. Small values of p produce 
low shock tension. However, if P is too small, the 
length of the adaptor becomes excessive. The solu¬ 
tion to the elimination of shock stress appears to be 
in the more careful detailed design of the adaptor. 


Design of Tapered Contour for 
Adaptor 


Because of certain aspects of the tooling problem, 
the designs for the 57/40-mm tapered-bore gun em¬ 
bodied a straight conical transition. It is now clear 
that a continuously curving transition contour is 
much better adapted to deform the projectile with 
minimum tensile stress. If the transition curve is 
made tangent to the straight muzzle bore section, the 
acceleration of the skirt from its minimum velocity is 
over a finite distance, and there is no tensile shock 
stress. In particular, if the transition contour and 
that of the skirt are both circular arcs, a very favor¬ 
able condition obtains. 

Letting s represent distance measured along the 
axis of the transition section from the throat forward, 
d the overall length of the transition, L the length 
of the section of skirt instantaneously in contact with 
the bore, and B 0 the entering slope of the section, the 
instantaneous acceleration of the sliding velocity is 
given by equation (8), which is obtained by differ¬ 
entiating equation (1). 

5= F„[ - sin 0 - sin (0 + «)]g§. (8) 


With sufficient approximation, 


s dp _ /So ds _ v 
L " U d 9 ds " d 9 dt °’ 


'■"G -Hr] 

[ . „ d — s , . ( . „ d — s 

sm po —^-b sin l a + 


)]■ 


(9) 


( 10 ) 


In all practical situations P is small enough to per¬ 
mit the use of p for sin p with sufficient accuracy. 
Furthermore, for the geometrical arrangement speci¬ 
fied, a is approximately a constant m times p. Hence, 
letting a = mp and using the approximation 
sin(a + P) = (m + 1 )p, the tensile stress T s is rep¬ 
resented by equation (11). 

T _ LqVq 2 Po 2 8(2 + P)(d — s)s / 11 x 

gd 3 ‘ U} 

Its maximum value is given by equation (12) 

T m = (2 + »)—( 12 ) 

The effect of this change on shock stress is strik¬ 
ingly shown by considering the case of a 57/40-mm 
gun with a 4-in. long transition. If we let T s = T 0 
= 100,000 psi; Lo = 0.6 in.; 8 = 0.28; d = 4 in.; 
m = 9 (90 degree arc); p 0 = 0.17 radian—the situa¬ 
tion will apply to a projectile of the type developed 
and the limiting velocity is represented by equation 

(13)_ 

r --4 uew .l + ,) - 9 ' ooot ^ (,8) 

With an 8 in. long arcuate transition, the limiting 
velocity increases to about 25,000 fps. 

By this change, the safe velocity has been more 
than doubled for a 4-in. taper, which tests showed 
would not pass a projectile at a velocity of 4,500 fps. 
The maximum safe velocity of an 8-in. taper having 
a continuously curving transition contour would thus 
be above the theoretical limit attainable with a gun 
of infinite length, which is assumed to be 20,000 fps 
(Section 3.5.2). Thus no practical limit to velocity 
appears to be imposed by the use of a tapered 
adaptor. In fact, furtherance of the development 
would present opportunities to reduce the length of 
adaptors while retaining a net gain in performance. 
There are good reasons to believe that, with the adop¬ 
tion of the smooth, horn shape for the transition, the 
length of the straight muzzle section may be ma¬ 
terially reduced. This is particularly true if means for 
controlling wear are adopted. There appears to be no 
reason why the length of this section should exceed 
twice the bearing length of the projectile, if the need 
for an excessive length allowance for wear is elim¬ 
inated. In the case of the 57/40-mm adaptor, reduc¬ 
tion of the working length from 30 in. to approx¬ 
imately 10 in. appears to be a reasonable possibility. 

Theoretically it is possible to design the transition 
to develop substantially constant tensile stress in the 


CONFIDENTIAL 










576 


TAPERED-BORE GUNS AND SKIRTED PROJECTILES 


skirt, throughout the travel period, but such an im¬ 
provement seems to have no practical importance. A 
decrease in horn length of only one-third could be 
effected, and this reduction is relatively insignificant 
in view of the more significant decrease in length re¬ 
sulting from the use of a circular contour. 


30 3 4 Muzzle Velocity and Velocity Limits 


When the muzzle velocity of the standard projec¬ 
tile is known the velocity of the corresponding sub¬ 
caliber projectile can be estimated with reasonable 
approximation by equation (14), in which 


V' = Fo 


(TFo + }TF p Y /2 

\W' + lWp/ 


(14) 


V' is the new velocity, F 0 the standard velocity, IF' 
the new weight, IF 0 the standard weight and W p the 
weight of powder. In the case of skirted projectiles, 
the preliminary estimates for a new design can be 
made from equation (15). 




1+ 4 


1 w? 


1/2 


Wo 


1 + p 


s / 4 K? 

Wo 


(15) 


Here p is the ratio of new caliber to standard caliber, 
8' is the mean density of the new projectile ma¬ 
terial, and S 0 the standard density. 

For estimating purposes, these approximations 
appear to be quite satisfactory for velocities up to 
5,000 fps in spite of a substantial increase in the 
kinetic energy imparted to the powder gases at these 
velocities. In the case of the 57/40-mm development, 
the standard projectile weight was 6 lb and its veloc¬ 
ity was 3,000 fps. The final design of carbide-cored, 
skirted projectile weighed 53 oz and attained a Veloc¬ 
ity of 4,100 fps when the powder was adjusted to give 
a maximum pressure equal to the standard pressure. 
This measured velocity compares well with a velocity 
of 4,050 fps calculated by equation (14). A skirted 
solid steel shot attained a velocity of 5,200 fps, when 
fired with a powder charge of 45 oz at the rated max¬ 
imum pressure for the gun (52,000 psi). This com¬ 
pares with 5,100 fps calculated by equation (15). 

For more exact computations, the methods out¬ 
lined in Chapter 3 are applicable. According to the 
analysis given there, the theoretical limiting velocity 
for an ideal, infinitely long gun is about 36,000 fps. If 
friction and heat losses are taken into consideration, 
with a high ratio of charge to projectile mass the 


limiting muzzle velocity is more likely to be of the 
order of 20,000 fps (Section 3.5.2). Tests with an 
8-mm, 125-caliber gun, carried out several years ago 
by Langweiler, 26 498 indicated a limit of about 14,000 
fps. Since, in such a small gun, heat and friction 
losses would be relatively high, the attainment of a 
somewhat greater velocity would be expected from 
the use of larger guns and guns in which is used some 
means of reducing friction, such as pre-engraved pro¬ 
jectiles (Section 27.3). 

The length of the 57/40-mm tapered-bore gun is 
50 calibers (with respect to the larger diameter). In 
spite of the fact that Langweiler’s gun was a little 
over twice as long (125 calibers, as compared with 
50) and that finely pulverized powder was used for 
its tests, velocities obtained with the tapered-bore 
gun firing skirted projectiles correspond to Lang¬ 
weiler’s results to within a few per cent. The charge- 
to-mass ratio ranged from 0.85 to 44. For a ratio of 10, 
the velocity is about 8,000 fps, which is 60 per cent of 
the indicated maximum of 14,000 fps. With higher 
ratios of charge to projectile, the velocity gain is 
slight; for a change in ratio from 10 to 20, the velocity 
increases by about 800 fps. It appears reasonable, 
therefore, to consider a velocity of 8,000 fps as repre¬ 
senting rather closely the practical limit to velocity 
attainable with present powders. Minor increases 
may be obtained by an increase in gun length, by 
recourse to measures designed to reduce friction, and 
by the development of higher potential propellants. 

In comparison with Langweiler’s gun, it is apparent 
that the lower efficiency of the tapered-bore gun, 
which results from its shorter barrel length and the 
relatively high friction of the skirted projectile, is 
compensated for by its lower heat loss. Analyses of 
deformation stresses indicate that no difficulty should 
be experienced in the construction of tapered-bore 
guns and skirted projectiles which would attain a 
velocity of about 8,000 fps. 

30-3-3 Venting to Suppress Gas Pressure 
Shock Wave at Base of Taper 

During early tests of the 57/40-mm gun, frequent 
difficulty was experienced with collapse and burning 
of powder-case necks. Since this prevented re-use of 
the cases, a remedy was sought. It was found that the 
phenomenon was generally absent when the projec¬ 
tiles were fired with the adaptor off, and that for each 
adaptor there was a velocity below which case failure 
did not occur. It was further found that the velocity 


CONFIDENTIAL 






DESIGN FOR TAPERED GUNS AND SKIRTED PROJECTILES 


577 


(or powder pressure) at which failure occurred was 
lower the shorter the taper. A typical burned case is 
shown in Figure 6. 

These observations led to the conclusion that col¬ 
lapse of the cases was caused by a shock wave set up 
when the projectile reached the base of the taper. 
This wave was a wave of pressure traveling toward 
the breech. Upon reaching the neck of the powder 
case it encountered an expansion from the bore di¬ 
ameter to the chamber diameter and experienced a 
partial reversal. If the amplitude of the shock wave 
was high enough, this reversal resulted in a pressure 
drop in the throat so that gases trapped between the 
neck and the bore caused collapse of the neck. Sub¬ 
sequent rapid escape of the gas in the chamber caused 
destruction by burning. British investigators who had 
encountered the same difficulty assumed that it was 
caused by the compression of air and gas between the 
skirts of the projectile. 402 

Further consideration of the situation led to a 
realization that the shock wave started at the adaptor 
could conceivably attain such amplitude as possibly 
to exceed the strength of the muzzle section of the gun. 
This leads to the question as to how much the ampli¬ 
tude of the shock wave may be. 

When a compressible fluid moving with relative 
velocity Vo strikes an obstacle, the stagnation pres¬ 
sure P s is given by equation (16), 

P a = , (16) 

in which P 0 is the initial pressure, K the gas constant, 
and N m , the Mach number corresponding to the rela¬ 
tive velocity. 

With the tapered-bore gun, the velocity of sound 
C in the gas at the base of the projectile as the projec¬ 
tile enters the straight muzzle bore is about 3,000 fps. 
The projectile velocity may be assumed to be 4,200 
fps. N m is therefore of the order of 1.4. A limiting 
value of P s , for the case of a zero-length taper, is 
given by the relations expressed in equation (17). 



The pressure Po at the base of the projectile is 
roughly 12,000 psi so the shock pressure will be 
around P 3 = 1.4 X 12,000 = 17,000 psi. This figure 


agrees reasonably well with a measured pressure of 
about 20,000 psi. 

To soften this shock wave, it is necessary to “bleed 
off’ 7 gas through vept holes at the base of the taper. 
The ideal way would be to provide vents having a 
capacity equal to the differential of projectile dis- 



Figure 6. On the right, type of cartridge case failure 
experienced when used with unvented tapered-bore gun 
tube. Undamaged case on left. (This figure has appeared 
as Figure 5 in NDRC Report No. A-456.) 


placement between the major and minor calibers. 
However, tests showed case failure occurred only 
above a critical velocity which was different for each 
taper. Only partial venting is therefore necessary. 

In the case of the 57/40-mm projectile at a velocity 
of 4,200 fps, the differential displacement is 58 cu ft/ 
sec, the density of the gas is about 7 lb per cu ft, and 
the mass displacement is therefore 400 lb per sec. The 
flow through a sharp edged orifice is given by 


Trr 0.53aP 
1 -l VT~’ 


(18) 


CONFIDENTIAL 







578 


TAPERED-BORE GUNS AND SKIRTED PROJECTILES 


in which W is mass, in lb/sec; a the orifice area, in 
sq in.; P the pressure, in lb/in. 2 ; and T the gas 
temperature, in degrees Fahrenheit on the absolute 
temperature scale. In the present instance P is about 
12,000 psi, T about 4000 F and, W, 400 lb/sec. 
Hence, the required vent area a is 4.2 sq in. 

It is interesting to note that in practice about 3 sq 
in. of vent holes were found necessary to prevent case 
failure. Four square inches was the area provided in 
the final designs. 

It was found that venting caused numerous gas- 
erosion grooves and gouges in the bore around the 
vent holes. These appeared to contribute to skirt 
breakage. The condition was rectified in later designs 
by machining a groove around the inside of the barrel 
just before the start of the taper. The vent holes were 
terminated in this groove. 

Figure 7 shows that nearly equal volumes of pow¬ 
der gas were discharged through the vents and from 
the muzzle of the gun. 



Figure 7. Effect of vents in tapered-bore gun tube on 
muzzle blast. (This figure has appeared as Figure 7 
in NDRC Report No. A-456.) 


30 4 GUN DEVELOPMENT TESTS 
30,4,1 General Testing 

In order to choose the optimum length of taper for 
deforming a two-skirted projectile, eight different 
tubes were designed and built. All had an overall 
length of 115 in. and a 40-mm section length of 7 in. 
In machining these tubes use was made of a thread 
collocating gauge designed for the purpose, 11 by 
means of which it is possible to locate the axial plane 
of a transverse center line across the threaded end of 


a tube such that the axial plane of this line bears a 
desired relation to the axial plane of a point on the 
face of the thread. 

The rate of taper used in changing the bore diam¬ 
eter from 57 mm to 40 mm varied from 34 in. per ft 
to 2 in., giving taper lengths of 32, 16, 8, and 4 in. 
One gun of each taper was rifled its full length and 
one of each left smooth in the taper. All of these tubes 
were rifled in the 57-mm section with 12 helical 
grooves having a pitch of 25 calibers. The selection of 
this rifling pitch was based on calculations. Later ex¬ 
perience with another tube showed that a pitch of 
17.5 calibers was too great. Although in general, a 25- 
caliber pitch performed satisfactorily, it was not used 
in the final design because the standard pitch (30 cal¬ 
ibers) was found to give the carbide-cored skirted 
projectile enough angular velocity for more than suf¬ 
ficient flight stability. 

For the smooth-taper barrels, skirt breakage was 
observed at a limiting maximum pressure of between 
25,000 and 37,000 psi copper, for taper lengths of 4 
and 8 in. Fired from the rifled, 16-in. taper barrel at 
copper pressures ranging from 46,000 to 52,000 psi, 
six shots remained intact. Satisfactory results were 
obtained with both the rifled and smooth-bored 32- 
in. taper, in tests at pressures as high as 40,000 psi. 

At the outset of the development, consideration of 
the motion of the projectile down the bore led to the 
conclusion that if the tapered and muzzle sections 
could be made short enough they could be left un¬ 
rifled, with negligible loss of spin and consequent sat¬ 
isfactory ballistic performance. Both ballistic tests 
and measurement of the pitch of scratches made on 
the smooth bore of various tubes showed that in all, 
except possibly the tube with the 32-in. long taper, 
the loss of spin was truly negligible. 

With the need for rifling in the tapered section thus 
eliminated, the last pretense of the necessity of single¬ 
piece construction was removed. Experience with the 
single-piece tapered-bore guns showed that the use of 
short smooth-bored tapers was practicable without 
sacrifice of performance. It therefore became obvious 
that the only practical construction for a tapered-bore 
gun was the use of a standard straight-bore gun with 
a smooth tapered-bore adaptor. This adaptor should 
differ from the Janacek adaptor in having a short, 
straight section of bore at the muzzle. It should also 
differ in having a curved transition section in place 
of a straight sided cone. Work was then concentrated 
on the development of a suitable, properly secured 
adaptor. 


CONFIDENTIAL 






GUN DEVELOPMENT TESTS 


579 


30 4 2 Testing of Adaptors 

The lightweight adaptor shown in Figure 2, weighs 
45 lb. Several experimental types were made to de¬ 
termine the optimum lengths of taper and 40-mm 
straight section. These results more or less substan¬ 
tiated the previous work on the one-piece tapered gun 
tubes. The short 4-in. taper caused breakups and was 
itself badly eroded after nine rounds. Tapers of 8, 16, 
and 32 in. in length performed satisfactorily with all- 
steel projectiles. The heavier tungsten carbide-cored 
projectiles were found to perform satisfactorily with 
the 8-in. taper. An 8-in. taper was therefore consid¬ 
ered to be the shortest practical length. 

Wear in the adaptors was fairly rapid. Adaptors 
were manufactured with a 40-mm bore section as 
short as 8 in. and as long as 22 in. The 21-in. length 
was decided upon not only because of the greater 
accuracy resulting from it, in comparison with the 
short lengths, but also because the life of the adaptor 
was increased. As can be seen in Figure 8, the wear on 
the adaptors was greatest near the end of the taper 
and decreased in the direction of the muzzle. In effect, 
the 40-mm section after some use assumed the form of 
a slow taper. Hence, the longer the 40-mm bore can be 
made, the greater the number of shots which can be 
fired while maintaining complete deformation of the 
projectile. 

Strain gauge measurements were made to deter¬ 
mine the safety factor which could be expected from 
the adaptor. In one case the measurements of cir¬ 
cumferential stress showed a total of 60,000 psi at the 
beginning of the taper, 44,000 psi at the end of the 
taper, and 49,000 psi in the middle of the straight 
section. These stresses all had the same general char¬ 
acteristics. There was a short period of high stress 
which lasted while the skirts were passing the gauges 
and then a relatively long period of stress caused by 
pressure decay in the gases. Skirt pressure stress 
started out at about 12,000 psi at the beginning of the 
taper, increased to 16,500 psi at the end of the taper, 
and jumped to 22,000 psi in the middle of the straight 
section. The gas pressure, on the other hand, dropped 
from 48,000 psi at the beginning of the taper to 27,500 
psi at the end of the taper, and 27,000 psi in the 
middle of the 40-mm straight section. These stresses 
are not excessive in consideration of the fact that the 
elastic limit of the adaptor steel is 110,000 psi. 

When these adaptors were attached to the stand¬ 
ard 57-mm Canadian Mk III gun of 87-in. length, 
the assembly was “muzzle heavy.” In order to pro¬ 


duce a balanced gun, the Canadian Mk III was 
turned down to the outside dimensions of a standard 
U.S. Mk V 57-mm gun. 



0 10 20 30 40 50 60 70 80 90 

ROUNDS FIRED 

Figure 8. Wear on tapered-bore adaptor No. 8 as a 

function of rounds fired. (This figure has appeared as 

Figure 16 in NDRC Report No. A-456.) 

30 4 3 Gun Wear Tests 

The wear in the guns tested was fairly high both in 
the adaptor and in the main bore at the origin of rifl¬ 
ing. This was to be expected. In the first place, the 
powder pressure exerts a substantial radial bearing 
component against the exposed area of the skirt. In 
the second place, the skirt is made of low carbon steel 
and it is a known fact soft steel bearing on steel under 
heavy bearing pressure produces a high rate of wear. 
When this frictional wear is added to the normal 
erosion wear at the origin of rifling, the observed high 
rate is not surprising. 

An attempt was made to silver-solder a pressed 
and sintered carbonyl iron band to the rear skirt, but 
the design of this application was not satisfactory. 

With the skirted-type projectile, wear occurred 
both in the grooves and in the lands. Projectile 
breakups noticed in worn tubes are probably caused 
by dislocation of the loosely fitting projectile in the 
enlarged section near the breech end at the moment 
of firing. Some gun tube data for wear on the lands 
are given in Figure 9. The fact that the muzzle wear 
is relatively slow compared with main tube wear cor¬ 
relates nicely with the fact that tube wear dies out 
much more rapidly along the bore than does the pow¬ 
der pressure which controls the skirt bearing pressure. 

During the early travel gas pressure is high, so the 
bearing pressure is high. Furthermore, at low speeds 
the coefficient of friction is high. As the speed builds 
up a surface film on the projectile apparently melts, 


CONFIDENTIAL 





























580 


TAPERED-BORE GUNS AND SKIRTED PROJECTILES 


and wear on the bore drops off rapidly. In this type 
of gun, the increased wear near the origin of rifling, 
can be explained by the fact that the reduced gas 
erosion, (resulting from a shorter time of travel) is 
offset by the increased mechanical friction in this 
region. Evidence of low friction at high velocity 
derives in part from the fact that the loss of spin in 
the taper is substantially less than can be obtained 
from calculations based on coefficients of friction at 
normal velocities. 



Figure 9. Wear on 57/40-mm gun tube No. 16 after 
80 rounds. (This figure has appeared as Figure 15 in 
NDRC Report No. A-456.) 


30 4 4 Attempts to Reduce Wear 

Various methods were attempted to increase the 
wear resistance of the adaptors by surface treatment, 
but these were, in general, unsuccessful or inconclu¬ 
sive. The first adaptor was made of NE 8713 steel, 
and gas carburized to a hardness of Rockwell C63 on 
the outer edge of the case. Cross sectioning andmetal- 
lographic examination of the muzzle end after six 
rounds had been fired through this adaptor showed 
surface cracks to a depth of approximately 0.035 in. 

On the basis of these results, later designs called for 
heat treating NE 8749 steel and drawing to Rockwell 
C30. These were more successful, although in the 
test of one adaptor, slight checking at the muzzle was 
observed after 103 rounds had been fired. No check¬ 
ing or scuffing was noticeable on any of the other 
adaptors heat treated in this manner. 

Another was chromium-plated to a thickness of 
0.003 in. at the National Bureau of Standards. After 
40 rounds no measurable wear was observed with 
gauges. However, at a distance of about 7 in. from the 
beginning of the taper there was a region where the 
surface was checked or scuffed. Also, in the straight 
section about an inch from the end of the taper, the 
chromium plate appeared to have flaked off of the 


surface. Chromium plate with its low coefficient of 
friction appears to offer good possibilities of increas¬ 
ing the life of these muzzle adaptors, if flaking can be 
controlled. 

Wear tests were made with a stellite liner (Chapter 
19) in a 57-mm muzzle adaptor. Twenty-six rounds of 
skirted projectiles, followed by 20 rounds of standard 
57-mm test projectiles, were fired through this liner 
with no apparent wear. There was a slight cracking 
observed. This was presumably caused, by the liner 
being out of round and not fitting tightly at all points 
in the adaptor. It is still to be hoped that a gun of the 
Janacek type with a stellite liner extending 12 to 15 
calibers forward from the origin of rifling, and with a 
stellite-lined muzzle adaptor would have a long life. 

30 5 PROJECTILE DEVELOPMENT TESTS 
30,51 Projectile Skirt Tests 

Rear Skirts 

A number of rounds was fired with steel-bodied 
projectiles to determine optimum rear-skirt thick¬ 
ness. The thicknesses tested varied between in. and 
% in. All showed good results at pressures up to 
40,000 psi in the unsegmented vented tube with a 
16-in. taper. These projectiles weighed about 33 oz. 
Other lightweight projectiles performed well with a 
%-in. thick rear skirt. However it was found, by tests 
with the tungsten carbide-cored type projectile (which 
weighed about 50 oz), that the minimum thickness 
that would stand up under firing was % in. 

Microflash pictures showed that the rear skirt had 
a slight tendency to balloon at high velocity under the 
action of the propulsive charge. A series of tests was 
undertaken to determine the effects on the skirts of 
high pressures. Slight flaring of the rear skirt, as ob¬ 
served from the size of the holes in cardboard screens, 
occurred between pressures of 45,000 and 55,000 psi 
(copper pressure). Breakage of the skirts was ob¬ 
served in one round at as low a pressure as 44,670 psi. 
No observable flaring was observed at a pressure as 
high as 50,900 psi. 

It can be concluded that the skirt-breakage limit 
for the recommended projectile skirt design is very 
close to the maximum pressure recommended by 
the Army for use with this type of gun. At pressures 
of 45,000 psi the muzzle velocity of the final design of 
the 53-oz armor-piercing projectile is estimated to be 
about 4,200 fps. 


CONFIDENTIAL 






















PROJECTILE DEVELOPMENT TESTS 


581 


Front Skirts 

The front skirt presented more of a problem. After 
trying a conical front skirt a number of experiments 
were made on skirts With a double curvature. None 
of these performed very well, and as far as could be 
seen from recovered pieces, the failures were in the 
front skirt. It was noted, however, that the thinner 
front skirts performed better than the thicker. One 
type had a tapered front skirt with a straight 

outline on the outside and a curvature at its base. 
However, wear measurements on this type indicated 
that the projectile was not deforming readily. It was 
therefore decided to revert to the uniformly curved- 
type front skirt. The thickness was standardized at 
Y s in. This type of front skirt has since performed 
quite satisfactorily. 

Skirt Position 

Not much attention was paid in earlier designs to 
the relative position of the front and rear skirts. Ac¬ 
tual practice indicated that this is a very important 
factor in the stability of the projectile both inside and 
outside the gun. The front skirt should be ahead of 
the center of gravity of the projectile at all times dur¬ 
ing deformation, to prevent the shell from careening 
inside the tube. In many of the later types of projec¬ 
tile, which were nose heavy as a result of the use of 
steel ballistic caps or high-density nose tips, an ab¬ 
normal amount of yaw was observed. 

A great deal of the wind resistance of the projectile 
is caused by the rear skirt and it, hence, has a large 
effect upon the center of pressure. 

The tendency has been in later designs, to keep the 
rear skirt as far back of the center of gravity as pos¬ 
sible by moving the core forward. In the recommended 
design type, the front skirt has been moved as far 
forward as possible without interfering with the con¬ 
tour of the ogive and the threaded section which re¬ 
ceives the windshield. 

Skirt Material Tests 

The first skirt material used was SAE 1020 steel. 
Though this was satisfactory as far as could be de¬ 
termined from the firings, it was decided to use SAE 
1120 steel because of its better machining qualities, 
a relative lack of ingot pattern, and its freedom from 
inclusions. SAE 3240, which had an original hardness 
of Rockwell B87-95, was found to be superior to both 


the other types with regard to breakage, but less satis¬ 
factory with respect to gun wear. During armor¬ 
piercing tests it was found that many projectiles 
made of SAE 1120 were breaking up under 45,000 psi 
maximum pressure. Steps were taken to strengthen 
the sheath in the thin region beneath the front skirt. 
Additional attempts were made to induction-harden 
the thin-walled section of the sheath. Tests showed 
that this method was not very satisfactory, and it was 
finally decided to resort to the heat treatment of the 
whole container to a hardness of Rockwell B84 to 
B90. Satisfactory results were obtained and this 
method was adopted. 

Windshield Design Tests 

The armor-plate penetration tests showed that 
the windshield design is extremely important at high 
angles of obliquity. There are apparently three pri¬ 
mary factors to be considered: 

1. To use an ogive with a long enough radius to 
prevent serious retardation in flight. 

2. To support the tungsten carbide core and pre¬ 
vent “shattering” upon impact. 

3. To prevent the projectile from “scooping” the 
armor plate at high angles of obliquity. 

Tests with one type projectile showed both “shat¬ 
ter” and excessive “scooping” for angles at which, 
judging from results of other high-velocity tungsten 
carbide-cored projectiles such as the 2-pounder 
Mk I “Littlejohns,” 154 there should have been per¬ 
foration with no “shatter.” By redesign of the wind¬ 
shield this performance was met. 

30 5 2 Test of Flight Characteristics 

There was some fear that the angular velocity pro¬ 
duced by a rifling of one turn in 30 calibers might not 
be enough to provide external flight stability for a 
skirted-type projectile. It was believed that frictional 
forces developed during passage through the taper 
section of the gun might retard the projectile consid¬ 
erably. 

The exterior-angular frequency F of a skirted pro¬ 
jectile can be obtained from yaw-card measurements 
of the rate of precession <t>' from equation (19), 

F - 0 , (19) 

in which A is the axial moment of inertia and B is the 
transverse moment of inertia. 


CONFIDENTIAL 



582 


TAPERED-BORE GUNS AND SKIRTED PROJECTILES 


The moments of inertia of the projectile were ob¬ 
tained from the measurements of the frequency of 
oscillation of the projectile, suspended from a wire, 
about the axis of each required moment of inertia 
The moment of inertia I about the suspension axis i g 
derived from equation (20). 


r T 2 na A 
“ 8t rh' 


( 20 ) 


in which h is the length of wire, y is the shear modulus 
of the wire material, a is the radius of the wire, and T 
is the measured time for a full oscillation. The ratio 
of moments of inertia is readily computed from equa¬ 
tion (21) 


-te> 


( 21 ) 


if the two times are measured with the projectile 
suspended from the same wire. By the same method 
the ratio of the axial moments of inertia before and 
after deformation can be obtained. 



Figure 10. Wave front of skirted projectile Xo. 451: 
Taken in flight; using microflash unit; velocity approx¬ 
imately 4,500 fps. (This figure has appeared as Figure 
21 in NDRC Report No. A-456.) 


These methods were employed to determine the 
following angular frequencies for one type projectile 
(No. 524): 

Angular Frequency 

1. New tube, no deforma¬ 
tion, no friction 67-in./turn (theoretical) 

2. New tube, deforma¬ 
tion, no friction 50-in./turn (theoretical) 

3. New tube (76 rounds) 

from measurements of <t>' 53-in./turn (observed) 

4. Old tube (189 rounds) 

from measurements of <f>' 67-in./turn (observed). 


In order to check these results an adaptor which had 
been used with a worn tube was cross-sectioned. 
Measurements made on the rifling imprints on this 
adaptor showed, as closely as could be determined, 
an angular frequency of one turn in 67 in. 

From these results it can be concluded that be¬ 
cause of the decrease in the moment of inertia after 
deformation, the angular velocity for this type pro¬ 
jectile should increase by about 35 per cent after 
passing through the taper section of the gun. Actually, 
observations showed an increase of about 25 per cent, 
the remaining 10 per cent being accounted for by 
frictional losses occurring during deformation. It can 
also be seen from these results that the projectile 
angular velocity decreases as the tube wear increases. 
Because of the increased angular velocity observed in 
a skirted projectile, it may be possible to tolerate 
more tube wear than in the standard gun and still 
maintain flight stability. 

Stability factors were calculated on the basis of 
data obtained from yaw cards spaced at 10-ft inter¬ 
vals along 200 ft of the firing range. From these the 
rate of precession, </>', and the length of one period 
of yaw, L, were measured. The stability factor S.F. 
is given by equation (22) : 287 


S.F. = ——- (22) 

(¥)- 

Calculations based on the results of the firings of the 
Type 524 projectile gave a stability factor of 1.8. In 
successive rounds, three types of projectiles which 
differed in the location of the tungsten carbide core 
were fired. The results are given in Table 2 of Chapter 
8. The core position of Type No. 1936 was used in the 
final design. 

Shadow pictures of projectiles in flight were ob¬ 
tained with the use of the Microflash equipment. The 
projectile velocities were approximately 4,400 fps. 
One of these pictures, shown in Figure 10, indicates 
very little distortion of the main wave front emanat¬ 
ing from the nose of the projectile. There are secondary 
waves from the region of the interface of the container 
and the ballistic cap and from the forward and back 
edges of the front and rear skirts. There is quite a bit 
of turbulence observable along the sides of the pro¬ 
jectile in the region in back of the front skirt. A large 
amount of turbulent wake is discernible in the rear of 
the projectile. From an analysis of the performance 


CONFIDENTIAL 





PROJECTILE DEVELOPMENT TESTS 


583 


data of the 57/40-mm skirted projectile, its ballistic 
coefficient was determined to be 1.0 for a form factor 
of 1.24. 

30.5.3 Windshield Testing 

Magnesium windshields- were tried in early projec¬ 
tile models, but this material was later discarded in 
favor of aluminum, mainly because of greater ease of 
machining. In firing tests against armor plate there 
was evidence that at high angles, a long aluminum 
windshield tends to swerve the core away from the 
normal to the plate, producing a “scoop” shot. This 
is treated more thoroughly in Section 30.6.3. 

Several projectiles were designed with steel bal¬ 
listic caps. They performed slightly better than the 
aluminum-capped type. 

Other projectiles were made with “soft steel” ballis¬ 
tic caps of various designs. There were usually only 
one or two projectiles manufactured of each of these 
designs and they were nearly all fired at 3-in plate at 
55 degrees. It is difficult to tell, therefore, which type 
of steel was more satisfactory. The main intent in 
using a steel ballistic cap is to help prevent the nose 
of the core from shattering upon impact. Thispurpose 
seemed to be realized. Compounded windshields of 
various materials were tried and were found to be the 
most efficient in the prevention of core shatter. 

Some projectiles had an elkonite nose tip and 
others were manufactured with elkonite nose pads. 
Elkonite is a powder metallurgy product composed 
of tungsten and copper. It has a density of 14.5 g/cc, 
approximately that of tungsten carbide, but unlike 
tungsten carbide, it is readily machinable. Its hard¬ 
ness ranges from 95 to 100 Rockwell B. Projectiles in 
which this material was placed ahead of the brittle 
tungsten carbide core were quite successful in pre¬ 
venting shatter at high angles of obliquity. Elkonite, 
however, is a relatively scarce and expensive material 
and was not used in the final design. The desired re¬ 
sults were obtained with a steel ballistic cap with an 
aluminum pad and tip. 

30 5 4 Testing of Cores 

General 

The core diameter was made 1.1 in. to conform to 
the core size of the Littlejohn 6-pounder skirted pro¬ 
jectiles and the British 6-pounder sabot-projectile. 
Some armor plate tests were made with projectiles 


which had finish-ground cores of the British design. 
Later unground cores were i^sed. These were similar 
to the ground one except for a 2.2-in. radius ogive on 
the radius instead qf the original 1.54-in. radius ogive. 
Projectiles were also manufactured with cores having 
a flat core nose tip. No significant variation in per¬ 
formance could be detected in tests against armor 
plate at high angles of obliquity. This seems to sub¬ 
stantiate the results of Division 2, NDRC. 156 

Physical Testing of Cores 

Cores of the following composition were used: 89- 
87% tungsten carbide, and 11 to 13% cobalt. 

Hardness measurements were made on 10 cylindri¬ 
cal ground cores and unground cylindrical cores. 
These, chosen at random from a larger lot, gave read¬ 
ings averaging 88.0 VDS Rockwell A. There was no 
significant difference in hardness between the ground 
and unground samples. Rockwell A88 converted into 
Vickers reading is 1350 DPH. This is considerably 
more than the minimum of 950 DPH required by 
British Ordnance for cores of these same specifications. 

The transverse rupture strengths of 14 tungsten 
carbide cores were determined. Eight unground cores 
measured had an average transverse rupture strength 
of 139,300 psi with a minimum of 111,700 psi and a 
maximum of 200,900 psi. Six unground cores measured 
had an average transverse rupture strength of 181,900 
psi with a minimum of 144,600 psi and a maximum 
of 215,700 psi. The minimum specification for British 
2-pounder SV Mk II unground cores is 156,800 psi. 
Two out of six unground cores failed to meet these 
specifications. Moreover, seven out of eight ground 
cores were found to have transverse rupture strengths 
below 156,800 psi. These results signify that the cores 
are measurably weakened when ground, probably due 
to thermal stresses impressed in the material by the 
operation. 

Transverse rupture-strength tests were made. 
Three tungsten carbide cores were used to apply the 
load to the core under test. The cores usually broke 
transversely into two main pieces. Considerable chip¬ 
ping was observable on the surface in the region of 
the fracture. 

There is no conclusive evidence gathered during 
the tests against armor plate to prove that the un¬ 
ground core performs noticeably better than the 
ground core. It is mainly on the basis of the trans¬ 
verse rupture strength tests that the unground core 
(Type 1499) is recommended for the final design. 


CONFIDENTIAL 



584 


TAPERED-BORE GUNS AND SKIRTED PROJECTILES 


see PERFORMANCE TESTING OF 
57/40-MM TAPERED-BORE GUN AND 
SKIRTED PROJECTILE 

30 6 1 General 

In an effort to stimulate interest of the Services in 
the potentialities of the tapered-bore gun, the ex¬ 
perimentation and testing were carried considerably 
beyond that necessary to establish the basis of design. 
In the course of this work the quality of performance 
of the gun was improved beyond the demands of 
practical necessity. The tests and the refinements 
developed during their conduct have shown results for 
the skirted projectile, in both accuracy and armor 
penetration, which are noticeably superior to either 
the sabot-projectile or the lightweight standard cali¬ 
ber projectile, of equal core size, when fired from the 
same gun. This advantage improves with the range. 



DEFLECTION IN INCHES 


Figure 11. Accuracy of skirted projectile J&L 860 at 
three velocities: Rounds 440 to 496. (This figure has 
appeared as Figure 31 in NDRC Report No. A-456.) 

30 6 2 Accuracy Tests 

Limitations on space available made it necessary 
to conduct most tests at a range of 100 yd. It is real¬ 
ized that a direct comparison between results at this 
range and at longer ranges is only approximate. How¬ 


ever, experience indicates that the initial yaw with a 
skirted projectile in a tapered bore is small. For this 
reason the extrapolation should be fairly satisfactory. 
Furthermore, the stability factor in the case of the 
armor-piercing projectile with a tungsten carbide 
core was rather high. It is therefore to be expected 
that the good results obtained should be fairly closely 
duplicated, under equally well-controlled conditions 
at longer range. 

The accuracy of the gun at 100 yd appears to de¬ 
pend on the type of projectile as well as the particular 
tube and adaptor which was used. Consequently, the 
accuracy chart shown in Figure 11 was based upon fir¬ 
ing in which these factors were constant. It was also 
found that the position of the shot groups relative to 
the bore sight depended upon the muzzle velocity. 
The scatter circles are drawn to include all shots of 
each of the three ranges of velocity. It can be noted 
that the accuracy of the projectile tested appears to 
improve as the velocity is increased. This may signify 
a more uniform deformation of the projectile at 
greater speeds or greater stability due to relatively 
less loss of spin in the adaptor. 

The probable error circle at 100 yd was computed 
for a projectile with a muzzle velocity of 4,200 to 
3,900 fps on the basis of the points shown in Figure 
11. This circle is shown in Figure 12. The probable 
error circle for the British 6-pounder sabot-projectile 
shown in the same figure, was obtained from nine 
shots fired in the standard 57-mm tube at muzzle ve¬ 
locities from 3,900 to 4,200 fps. Assuming a normal 
distribution, this circle represents the area into which 
50 per cent of the shots are expected to fall. If the dis¬ 
persion be assumed to be proportional to the range 
the radius of the probable error circle would be 7 in. 
at 1,000 yd for the skirted projectile, while it would 
be over 3 ft at this distance for the sabot-projectile. 
The superior accuracy of the skirted projectile is 
probably due, in part, to the very low value of initial 
yaw and to a somewhat higher stability resulting 
from a more favorable location of center of gravity. 
It is, of course, possible that the low accuracy of the 
sabot-projectile is due in part to disturbance of it in 
flight by the discarding elements. 

Firings were later conducted at Aberdeen Proving 
Ground. 540 The gun was mounted in a Scout car and 
firings were made against a 12 ft by 12 ft target at a 
range of 1,717.63 ft. The average instrumental velocity 
for 20 rounds was 3,950 fps. 

The radius of the probable error circle under these 
conditions was 12 in. at the above mentioned range 


CONFIDENTIAL 





















TESTING OF 57/40-MM GUN AND PROJECTILE 


585 


or 21 in. at 1,000 yd, assuming, again, that the trajec¬ 
tory is a straight line. 

This dispersion was somewhat greater than was to 
be expected on the basis of firings at the Jones and 
Lamson range. The discrepancy may be caused either 
by the two different types of gun mountings used, or 
by the fact that an unground core was supplied with 
the projectile type fired at the Proving Ground, while 
a ground core was used in the other projectile. 


30 6 3 Armor Penetration Tests 

Preliminary Tests 

Aberdeen Proving Ground was supplied with the 
tubes and adaptors and with two types of skirted 
projectiles. These were used to conduct a series of 
tests 560 to determine the performance of the tapered- 
bore gun against 4-and 6-in. armor plate at 30 degrees 
obliquity, and 3-in. armor at 55 degrees. Members of 
the Princeton University Station, Division 2, NDRC, 
analyzed these firings and compared the results with 
those of other tungsten carbide-cored projectiles. 154 
The performance of the skirted projectile was much 
worse than was to be expected. It was concluded at 
this time, after a study of the results, that the un¬ 
satisfactory behavior was probably due to a disinte¬ 
gration of the tungsten carbide core at the moment of 
impact. The Jones and Lamson Research Department 
in cooperation with members of Division 2, NDRC, 
undertook a program of armor-plate tests to deter¬ 
mine the ballistic limit of the skirted projectile, and 
to devise an improved design in which the core would 
not shatter upon impact. 


Design Changes to Improve Penetration 

A number of design changes was made to attempt 
to prevent core shatter. Of these the following are 
most noteworthy. 

1. An aluminum nose pad was added. This did not 
prevent shatter when used alone, but seemed to be 
advantageous in conjunction with a steel ballistic cap. 

2. A steel windshield was substituted for the alumi¬ 
num windshield. 

3. An unground core was substituted for the ground 
core. 

In addition to the problem of core shatter it was 
noticed that the length of the distance from the core 
nose to the tip of the ballistic cap had an important 
effect on the performance at high angles of obliquity. 


The plate results showed that the long aluminum bal¬ 
listic cap of the projectile was deflecting the core, 
before it hit the plate, to such an extent as to produce 
“scooping.” To counteract this effect the windshield 
was shortened and the steel tip replaced with an 
aluminum tip. This type of tip not only increases the 
stability inside of the gun by shifting the center of 
gravity backwards, but also provides a soft nose 
point which produces less swerving away from the 
normal to the armor plate at the moment of impact. 

Several other minor changes were made. A radius 
of curvature was added to the section under the front 
skirt to replace the sharp angle where container break¬ 
age was observed. In addition the container was hard¬ 
ened overall to Rockwell B84 to B90. Later tests 
showed that, with this treatment, the skirts would 
not break at 45,000 psi (copper pressure), the rated 
pressure of the gun. 


Ballistic Limits of Skirted Projectile and of 
British Sabot-Projectile 


The ballistic or perforation limit of any partic¬ 
ular type of projectile depends upon the following 



Figure 12. Probable error circles at 1000 yds for 
skirted projectile J&L 860 and British 6-pounder sabot- 
projectile: Velocity 4,020 ± 160 fps. (This figure has 
appeared as Figure 32 in NDRC Report No. A-456.) 


CONFIDENTIAL 













586 


TAPERED-BORE GUNS AND SKIRTED PROJECTILES 


three factors: 

1. The striking velocity of the projectile. 

2. The angle of obliquity of the armor plate. 

3. The thickness of the plate. 

Firings were made against 2-in., 3-in., 4-in., and 
6-in. homogeneous armor plate supplied by Aberdeen 
Proving Ground. Three levels of projectile velocity 
were used; 3,200 fps, 3,700 fps, and 4,150 fps. The 
plate was set at 5-degree intervals and an attempt was 
made to bracket the ballistic limit with two shots on 
either side at each of the velocity levels. From the 



0 10 20 30 40 50 60 

PLATE ANGLE IN DEGREES 


Figure 13. Ballistic limit as a function of plate thick¬ 
ness and plate angle. (This figure has appeared as 
Figure 35 in NDRC Report No. A-456.) 

size of the holes made when perforation occurred, 
observation of recovered material and from gamma- 
graphs and exographs it was evident that above cer¬ 
tain angles and velocities the skirted projectile com¬ 
pletely shattered on impact. The minimum velocity 
at which shatter occurred was, in general, between 
2,500 and 3,000 fps. It occurred at angles of obliquity 
which varied with the thickness of the armor plate. 
With 2-in. armor plate one type of skirted projectile 
shattered at angles roughly greater than 50 degrees; 


for 3-in. plate at angles greater than 35 degrees; for 
4-in. plate at angles greater than 25 degrees, and for 
6-in. plate at angles greater than 10 degrees. Forty 
British 6-pounder discarding sabots with 1.1-in. 
tungsten carbide cores were made available for firing 
at the Jones and Lamson range. Firings made at the 
same plates used for the skirted projectiles offered a 
means of comparing the relative performance of the 
three types. 

One type of skirted projectile and the British sabot- 
projectile performed much better than the other type 
at high angles of obliquity, mainly because of the 
absence of core shatter. A third type of skirted pro¬ 
jectile (No. 1949) which is identical to the final de¬ 
sign except for an insignificant detail in the container, 
penetrated 3 in. of armor at an obliquity angle of 55 
degrees. It has been noted that the performance of 
the skirted projectile is better than that of the sabot- 
projectile at low angles of obliquity. This is appar¬ 
ently due, in part, to the greater weight of parts of 
the sheath in relation to the weight of corresponding- 
parts of the British projectile. The weights of the two 
types of skirted projectile are approximately 48 and 
50 oz, respectively, while the weight of the sabot-pro¬ 
jectile is only 37 oz. The greater weight results in 
higher residual velocity at impact and in higher ef¬ 
fective mass at impact. However, as the angle of 
obliquity of the armor plate is increased the sabot- 
projectile gradually overcomes this initial disadvan¬ 
tage. Two reasons are advanced to explain this action. 
In the first place, the distance from the core nose 
to the point of the ballistic cap is only 1.00 in. in 
the sabot-projectile design, while it is 1.31 in. in the 
design of the skirted projectile. In the second place, 
the subcaliber sheath diameter of the sabot-projectile 
is only 1.45 in., as opposed to a sheath diameter of 
1.58 in. in the case of the skirted projectile. 

In Figure 13 the ballistic limit is shown as a func¬ 
tion of plate thickness and plate angle for three dif¬ 
ferent velocities. Curves for Type No. 1944 skirted 
projectile seem to converge at an angle of approxi¬ 
mately 70 degrees. Above this angle, the projectile 
would be expected to ricochet regardless of projec¬ 
tile velocity or plate thickness. The “ricocheting 
angle” for the sabot-projectile appears to be about 
65 degrees, which is reasonable in view of the fact 
that the projectile is somewhat lighter. The maximum 
thickness which Type No. 1944 would be expected to 
penetrate would be about 9.4 in. of armor at normal, 
while the sabot-projectile would be expected to pene¬ 
trate approximately 8 in. at normal. 167 Figures 14 


CONFIDENTIAL 






























TESTING OF 57/40-MM GUN AND PROJECTILE 


S87 



Figure 14. Perforation produced by projectile 1949 in 3-in. armor plate at an angle of 55-degree obliquity: Front of 
plate 3G; two-thirds actual size. (This figure has been taken from a Division 2 report.) 


and 15 illustrate the absence of core shatter from the 
firing of a Type No. 1949 skirted projectile at a ve¬ 
locity of 4,200 fps, and an angle of obliquity of 55 
degrees. 


in Section 9.2. These results are shown in Figure 
16 for the cases of 0, 30, and 50-degree angles of ob¬ 
liquity. Both a and the constant & varied with the 
impact angle according to data in the following table. 


De Marre Formula 


An attempt was made to fit the penetration results 
of the skirted projectile to the De Marre formula, 


(e\ 8 TFF 2 cos 2 0 
“W “ d 8 • 


(23) 


This gives the penetration, e (in.), of a projectile 
in terms of the striking velocity V (fps), angle of 
obliquity 6 degrees projectile weight W (lb), and pro¬ 
jectile diameter d (in.). This application is discussed 


Angle 

(degrees) 

Logio « 

P 

0 

5.88 

1.60 

30 

6.04 

1.45 

50 

6.19 

1.26 


For steel projectiles equation (25) was found 564 to 
be quite accurate with 0 = 1.43 and logio a = 6.15. 

The British find that with tungsten carbide cores, 
the constant logio a varies with the angle of attack 
as was observed in the case of the skirted projectile. 


CONFIDENTIAL 










588 


TAPERED-BORE GUNS AND SKIRTED PROJECTILES 


Using a core of 0.653-in. diameter and 0.42-lb weight 
against 80-mm and 100-mm homogeneous armor at 
striking velocities between 3,000 and 3,500 fps, the 
following values of logioa were observed; at 0 degree, 
log 10 a = 5.965; at 30 degrees, logio = 6.117 — 
0.017 (e/d — 1). The observations made at the J&L 
range appear to be in very good agreement with 
these results. Figure 16 shows that carbide cores will 
average about three times the penetrating power of 
steel cores. 

Perforation Range Comparisons 

Using the data supplied by Princeton University 
Station, Division 2,NDRC,and those of the British 301 


in conjunction with the results of the firings at the 
J&L range, the perforation range of the J&L 1944 
and the British sabot were determined. The retarda¬ 
tion of the J&L Type 1944 was assumed to be the 
same as the retardation of the J&L Type 860. The 
perforation-range curve of the sabot given by the 
British is for armor plate at 30-degree angle. In order 
to obtain the retardation of the sabot the following 
procedure was used. From the curves for the ballistic 
limit as a function of striking velocity and angle of 
obliquity for 3-in., 4-in., and 6-in. plate,the 30-degree 
line for the British sabot was drawn on Figure 17. As 
can be seen, this is in close agreement with the datum 
point taken from the perforation range curves for 6- 
pounder projectiles supplied by the British. Conse- 



Figure 15. Perforation produced by projectile 1949 in 3-in. armor plate at an angle of 55-degree obliquity: Rear of 
plate 3G; two-thirds actual size. (This figure has been taken from a Division 2 report.) 


CONFIDENTIAL 








TESTING OF 57/40-MM GUN AND PROJECTILE 


589 


quently, this line was used in conjunction with the 
British data to obtain the retardation curve for the 
British sabot. Using the retardation for the British 
sabot and the J&L projectiles, and the ballistic limit 
curves for 3-in. and 6-in. armor plate, the perforation 
range as a function of the obliquity angle was deter¬ 
mined for J&L projectile Type 1944 and the British 
sabot against these two thicknesses. These curves 
show plainly that while the British sabot is slightly 
more effective than the J&L 1944 at short ranges, it 
rapidly loses its advantage at distances greater than 
400 or 500 yd. This is even more strikingly apparent 
in the case of the 6-in. plate than in the case of the 
3-in. plate. There are apparently two reasons for this 
behavior. In the first place the subcaliber sabot, being 
only about 75 per cent as heavy as the J&L Type 
1944, loses velocity at a more rapid rate. In the sec¬ 
ond place, the greater amount of energy at the mo¬ 
ment of impact, due to the greater mass as well as 
the higher velocity, acts in favor of the J&L Type 
1944 projectile. 

Using the ballistic limit curves, it was possible to 
obtain the specific limit energies for the J&L pro¬ 
jectile, at various plate angles up to 55 degrees, as a 
function of the plate thickness in core calibers. In 
these calculations the weight and diameter of the 
core has been used. The remainder of the projectile 
was neglected. The average performance for each 



Figure 16. Determination of constants in DeMarre 
formula for tungsten carbide cores against armor plate 
at impact angles of 0 degree, 30 degrees, and 50 degrees, 
and comparison with values for steel cores. (This figure 
has appeared as Figure 44B in NDRC Report No. 
A-456.) 


angle was approximated by straight lines on a loga¬ 
rithmic plot. On the basis of this graph, predictions 
have been made of the ballistic limits, against 3- and 







• 

4 

,30* BRIT 
/ BRITI 

1 

ISH SABOT 
ISH DATUM 








’ BRITISH S 
/ BRITK 

SABOT 

5H DATUM 





V 

L,^ 

r ^ 


*J AND L 8i 
J AND L 1 

60 




1 




DATUM 




1 r *> 







"c 

S' 









3 4 5 6 7 

PLATE THICKNESS IN INCHES 


Figure 17. Ballistic limit as a function of striking velocity and plate thickness. (This figure has appeared as Figure 
44C in NDRC Report No. A-456.) 


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590 


TAPERED-BORE GUNS AND SKIRTED PROJECTILES 


6-in. plate, for a 90/63-mm skirted projectile, which 
had been scaled up from the 57/40-mm projectile 
dimensions. These calculations were made on the 
basis of a core weight of 6.04 lb and a core diameter of 
1.75 in. for the enlarged projectile. 

A prediction of the perforation range for a 90/63- 
mm projectile against 3-and 6-in. armor plate is shown 
in Figure 18. The retardation for this prediction 
was considered to be inversely proportional to the 
projectile weight. 



Figure 18. Perforation range as a function of angle 
of obliquity for J&L Type 944 and British 6-pounder 
sabot: Muzzle velocity, 4,150 fps. (This figure has 
appeared as Figure 44E in NDRC Report No. A-456.) 


30 7 SUGGESTIONS FOR FUTURE 
DEVELOPMENT 

Any plans for future development of tapered-bore 
guns and skirted projectiles should include attention 
to the following: 

1. Application of a thin layer of bronze or other 
bearing material on the bearing surfaces of the skirts 
of projectiles: This should greatly improve the wear 
and accuracy life of gun tubes and tapered adaptors 
for the observed wear therein is predominantly 
through friction. 

2. Application of wear resistant liners of such 
material as stellite in gun tubes and adaptors: 
With suitable bearing material coatings on projec¬ 
tiles and wear resistant surfaces in gun tubes and 
adaptors the life of tapered-bore guns fired at 
present pressures should equal or exceed the life of 
standard guns. 

3. Development of a suitable high-explosive type 
skirted projectile for antiaircraft use: The potential 
advantages of subcaliber projectiles in this service 
are tremendous. 

4. Detailed analysis of the shock stresses, plastic 
deformation, and friction effects in projecticle skirts: 
It appears most likely that improvements in maxi¬ 
mum velocity without breakage and reduction in 
length of tapered bore and straight section of adap¬ 
tors may result from such studies. 


CONFIDENTIAL 








































Chapter 31 

PRE-ENGRAVED PROJECTILE WITH CHROMIUM-PLATED BORE 

By Nicol H. Smith a 


311 INTRODUCTION 

I n the early work with the caliber .50 erosion¬ 
testing gun described in Section 11.2.1, it was soon 
realized that for hypervelocities the standard caliber 
.50 gun barrel and bullet were not satisfactory. The 
gilding metal jacket was too weak to withstand the 
rotational stresses due to the high velocity as long as 
the depth of the grooves was as small as it is in this 
barrel. This deficiency was partly overcome by using 
a banded projectile with a copper rotating band, but 
it needed to be excessively wide for satisfactory 
performance. 

In addition, there was little or no exact information 
concerning (1) the relative effect on gun erosion of 
wear produced by the engraving and rubbing of the 
projectile and (2) the influence of gas leakage (and 
obturation) on gun erosion. These questions could be 
answered by experimental firings of pre-engraved 
projectiles in the caliber .50 erosion-testing gun. With 
this end in view, work on pre-engraved projectiles 
(hereafter referred to as PE) was started at The 
Franklin Institute in January 1943. 

The use of PE projectiles makes possible the reduc¬ 
tion to a minimum of engraving stresses and bullet 
friction, properties inherent in projectiles which are 
not pre-engraved. Thus it is possible to eliminate 
these factors as contributing causes of gun erosion 
and isolate that which may be attributed to powder 
gases alone. 

Hence, the purpose of this investigation became 
threefold. 

1. To determine the effect of PE projectiles upon 
gun erosion, pressure, velocity, and accuracy; 

2. To examine the exterior ballistics of PE projec¬ 
tiles of different design; and 

3. To see how such projectiles can be adapted to 
firing in existing guns of different caliber. 

Early work on PE projectiles had been done in 
England and France, especially by Charbonnier dur- 

a Associate Director, The Franklin Institute, Philadelphia, 
Pa. 

b Sections 31.1 to 31.7 have been based on NDRC Report 
A-448 120 by the same author, to which reference should be 
made for further details. All the tables except No. 12 and all 
the figures except No. 19 have been taken from that report. 


ing World War I. Reference is made to the more 
complete report 120 for a historical sketch of PE pro¬ 
jectiles. They were used by the Germans in World 
War II for at least one gun, a 28-cm railway gun. 303 304 

312 CONCLUSIONS 

The use of PE projectiles has enabled one to de¬ 
termine the relative effect of the components affect¬ 
ing gun erosion. Results have shown that (1) the 
engraving of the rotating band and the abrasion 
produced by the friction of the bullet are important 
factors contributing to the erosion of both gun steel 
and chromium plate; (2) gas leakage is not a direct 
cause in initiating gun erosion; and (3) reduction of 
engraving stresses and bullet abrasion reduces the 
wear of gun steel and chromium-plated gun bores. 
The use of PE projectiles in determining the erosive¬ 
ness of different propellants is described in Section 
15.3. 

The reduction of bullet wear results in an increase 
in the velocity life of the gun. This increase is about 
twofold to fivefold in gun steel bores and twentyfold 
in chromium-plated bores. 

PE projectiles permit the use of high-strength steel 
bands, thereby resulting in a very material improve¬ 
ment in the accuracy of the projectile. 

Experiments with the caliber .50 Browning machine 
gun mechanism have shown that PE projectiles prop¬ 
erly indexed can be fired in a gun having a high cyclic 
rate of fire without hang-ups or any malfunction of 
the gun attributable to the PE projectile, as described 
in Section 28.4. 

The use of the steel-banded PE projectile chro¬ 
mium-plated bore combination makes it possible to 
obtain a hypervelocit}^ gun having a greater velocity- 
and accuracy-life. 

313 METHODS OF TESTING AND 
MEASUREMENT 

The caliber .50 erosion-testing gun (Section 11.2.1) 
was used as the tool in the preliminary tests of the 
PE bullets. In most tests double-base powder con¬ 
taining 20 per cent nitroglycerin was used. 


CONFIDENTIAL 


591 



592 


PRE-ENGRAVED PROJECTILE WITH CHROMIUM-PLATED BORE 



X 

< 

2 

O 

oj 


i 


The PE bullet shown in Figure 1 was adopted as 
the standard steel-banded PE projectile for test pur¬ 
poses. Conditions of loading were chosen in establish¬ 
ing the powder charge to give a maximum powder 
pressure of 56,000 to 58,000 psi (copper). The web of 
the powder was adjusted by mixing a “fast” powder 
(small web) with a complementary part of a “slow” 
powder (large web) so that a charge of 476 grains 
gave the desired pressure and a velocity of about 
3,700 fps in a 45-in. barrel. 

Velocity and pressure measurements were obtained 
at intervals to determine the velocity-life of the gun. 
Plug gauge and star gauge measurements were made 
at intervals to determine the distribution and prog¬ 
ress of erosion of the lands and grooves. Accuracy 
measurements were taken at 100 ft to determine the 
accuracy-life of the gun. 

The gun was considered to have failed when: (1) 
the velocity had dropped 200 fps, or (2) the mean 
radius of dispersion at 100 ft had increased to three 
times its initial value or keyholing bullets were ob¬ 
served. 

314 CALIBER .50 EROSION-TESTING 
GUN EXPERIMENTS 

31,4-1 Introduction 

PE projectiles were fired in gun barrels having a 
steel surface and a chromium-plated surface. In both 
cases the effect of PE projectiles on the following was 
studied. 

1. Progress and distribution of land and groove 
erosion. 

2. Pressure and velocity performance. 

3. Accuracy performance. 


In addition, using PE projectiles as a tool, the 
effects of the following factors on the erosion of gun 
steel and of chromium plate were studied. 

1. Gas leakage and obturation. 

2. Type of projectile. 

3. Thickness of chromium plate. 

The results and conclusions of these tests are given 
in the following sections. 

31,4,2 Effect of Gas Leakage and 

Improved Obturation on the Erosion 
of Gun Steel 

Since the standard PE projectile has a clearance of 
0.002 to 0.003 in. across the lands and grooves in the 
bore, the first question which is presented is: “What 
effect will this gas leakage have on erosion?” To an¬ 
swer this question, the following tests were fired. 

1. Standard PE projectiles for 0.005-in. rifling 
which gave a calculated leakage area of 0.0028 sq in. 

2. PE projectiles having two slots 0.020-in. deep 
milled diametrically opposite, which gave a calculated 
leakage area of 0.0078 sq in. 

3. PE projectiles having two slots 0.030-in. deep 
milled diametrically opposite which gave a calculated 
leakage area of 0.0103 sq in. 

4. Standard PE projectiles with cellulose obturat¬ 
ing wads saturated with liquid fluorocarbon. 

5. Standard PE projectiles with solid fluorocarbon 
obturating wads. 

Liquid fluorocarbon is a very stable organic com¬ 
pound (polymerized polyfluorethylene derivative) 
which under the conditions of firing does not foul the 
bore by its decomposition products. This material has 
the consistency of heavy molasses. Cellulose wads 
were saturated with the heated liquid and attached 
to the base of the PE projectile. 


CONFIDENTIAL 


































CALIBER .50 EROSION-TESTING GUN EXPERIMENTS 


593 


The solid fluorocarbon was the same basic material 
but polymerized further to the solid condition. Wads 
were cut to fit the neck of the cartridge case and 
attached to the base of the projectiles. The fluoro¬ 
carbons were obtained from Division 8, NDRC. 

When making the firing test the successive rounds 
were so chambered that the slots in the projectile 
were progressively rotated one barrel groove per 
round, thus insuring equal presentation of the slot 
around the circumference of the bore. 

In order to keep the pressure and powder charge 




O 20 40 60 80 100 120 140 160 180 200 

ROUNDS FIRED 


LEAKAGE AREA 

1 .0028 SQ IN. 

2 .0078 SQ IN. 

3 .0103 SQ IN. 


constant at 56,000 to 58,000 psi (copper) and 476 
grains, respectively, it was necessary to increase the 
percentage of fast powder as the leakage area was 
increased. 

A summary of the firing conditions and the ob¬ 
served increase in land and groove diameters at 0.5 
in. beyond the origin of rifling are given in Table 1. 
The latter data are shown graphically in Figures 2 
and 3. 




0 20 40 60 80 100 120 140 160 180 200 


ROUNDS FIRED 


LEAKAGE AREA 

1 0.002 8 SQ IN. 

2 0.0103 SQ IN. 

3 0 

4 0 


OBTURATOR 

NONE 

NONE 

CELLULOSE WAD 
FLUOROCARBON WAD 


Figure 2. Effect of projectile clearance on land and 
groove erosion, for PE projectiles fired in the caliber 
.50 erosion-testing gun. 


Figure 3. Effect of obturation of PE projectiles on 
land and groove erosion of a steel barrel in the caliber 
.50 erosion-testing gun. 


Table 1. Firing conditions and land and groove erosion for gas leakage tests with caliber .50 pre-engraved projectiles 
after 35, 70, 105, 140, and 180 rounds. 


Test 

Leakage area 
(in.) 

Pressure* 

(psi) 

Velocity 

(fps) 


Increase 

in land diameter 


Increase 

in groove 

diameter 


35 

A L in in. X 
70 105 

10~ 3 

140 

180 

35 

AG 

70 

in in. X 
105 

lO- 3 

140 

180 

L1(F4) 

0.0028 

57,200 

3,753 

5.6 

9.6 

14.4 

18.7 


4.2 

6.5 

9.3 

12.0 


F(F4) 

0.0078 

54,900 

3,630 

4.4 

8.2 

11.6 

15.0 

19.3 

4.0 

6.3 

8.8 

11.1 

14.1 

F(F5) 

0.0103 

57,500 

3,678 

2.8 

5.7 

9.1 

12.5 

15.8 

2.3 

4.5 

6.6 

8.8 

11.7 

F(F3) 

Of 

56,700 

3,705 

4.0 

7.6 

11.2 

14.8 

18.8 

2.6 

6.0 

8.2 

10.5 

13.5 

F(F6) 

0§ 

55,800 

3,701 

5.0 

9.2 

13.0 

16.7 

21.0 

4.0 

6.4 

9.9 

12.5 

15.4 


* All tests fired with 476 grains double-base powder of two granulations. 
fCellulose wads. 

§Fluorocarbon wads. 


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594 


PRE-ENGRAVED PROJECTILE WITH CHROMIUM-PLATED BORE 


These results show that the increase in land and 
groove diameter is least for the greatest gas leakage 
area. The data for A L and AG fall in the following 
order: 

AL 0.0028 in. 2 > 0.0078 in. 2 > 0.0103 in. 2 leakage 
area. 

AG 0.0028 in. 2 > 0.0078 in. 2 > 0.0103 in. 2 leakage 
area. 

AL Solid fluorocarbon wad > cellulose-liquid 
fluorocarbon wad > 0.0103 in. 2 leakage area. 

AG Solid fluorocarbon wad > cellulose-liquid 
fluorocarbon wad > 0.0103 in. 2 leakage area. 

The land erosion observed with both types of ob¬ 
turating wads was slightly less than observed with 
the standard PE projectile (0.0078 sq in. leakage 
area). However, in both cases the erosion was greater 
than observed with the projectiles having the greatest 
leakage area (0.0103 sq in. leakage area). 

These test data show that for a caliber .50 PE 
projectile the clearance between the projectile and the 
bore neither initiates nor accelerates gun erosion. In 
fact, with increasing leakage area it was found that 
the measured land and groove erosion decreased. It is 
believed that this reduction in erosion is caused by 
the increasing percentage of fast powder. An increase 
in the percentage of fast powder would result in a 
shorter duration of maximum pressure and hence 
maximum temperature. 

Since the PE projectile is free moving in the bore 
of the gun, any gas leakage past the projectile would 
only expose one particular area of the bore to the 
blast of hot gas for a very short time. This is contrary 
to the situation with the regular engraving-type 
projectile, which is momentarily slowed up in its 
travel during the engraving process. Any gas leakage 
past the engraving bullet would accelerate gun erosion 
because of the “blow torch” action discussed in Sec¬ 
tion 5.4.4 of the gases for a much longer time on a 
localized area of the bore. 


Serious effects due to gas leakage with PE projec¬ 
tiles might be experienced in a larger caliber gun, 
where the time required to start the projectile mov¬ 
ing might be unusually long. In this case the product 
of duration (time) X heat (cal/cm 2 ) might be 
great enough to increase erosion. In this case it would 
be necessary to use some practical method for obtu¬ 
rating the PE projectile. 


314 3 Effect of Pre-Engraving on 

Pressure and Velocity Performance 

Starting Resistance and Powder Pressure 

The manner in which a given propellant burns, 
which determines the maximum powder pressure, is 
dependent mainly upon the resistance to starting the 
projectile. In a conventional type of gun using banded 
ammunition, this starting resistance can be attributed 
to: (1) force to overcome the inertia of the projectile, 
(2) force to unseat bullet from case, and (3) force to 
engrave the rotating band. 

As the starting resistance decreases, the burning 
characteristics of the powder are so changed as to 
affect seriously the maximum powder pressure. Using 
PE projectiles the force to engrave is completely 
eliminated, with the result that the powder pressure 
obtained with the same granulation of propellant as 
used for an engraving bullet is very much less. 

The effect of decreasing the force to engrave is 
clearly shown in Table 2. The force to engrave was 
gradually decreased by reducing the band diameters 
and finally eliminated by pre-engraving the projec¬ 
tiles. A pressure drop of 17,900 psi and a velocity 
drop of 292 fps were observed. 

In order to keep the pressure at the level observed 
with an engraving-type bullet, keeping the powder 
charge the same, it is necessary to use a powder with 


Table 2. Effect of engraving force on pressure and velocity of caliber .50 projectiles. 


Band 

dia. 

(in.) 

Land 

dia. 

(in.) 

Band/Land 

interference 

(in.) 

Pressure* 

(psi) 

Pressure 

change 

Velocity 

(fps) 

Velocity 

change 

0.510 

0.490 

0.020 

54,000 


3,452 


0.509 

0.490 

0.019 

54,200 

+ 200 

3,405 

- 47 

0.508 

0.490 

0.018 

52,700 

- 1,300 

3,380 

- 72 

0.503 

0.490 

0.013 

49,100 

- 4,900 

3,330 

-122 

0.498 

0.490 

0.008 

44,800 

- 9,200 

3,280 

-172 

0.493 

0.490 

0.003 

41,600 

-12,400 

3,233 

-219 

(PE) 

0.490 


36,100 

-17,900 

3,160 

-292 


* All firings made with a charge of 465 grains IMR powder, web thickness 0.0338 in. 


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CALIBER .50 EROSION-TESTING GUN EXPERIMENTS 


595 


a smaller web thickness (“faster” powder). In the 
experiments just referred to, the same charge of a 
powder having a web thickness of 0.0285 in. gave the 
same pressure and a greater velocity than the engrav¬ 
ing-type bullet. The relationship between type of 
projectile, web thickness, pressure and velocity is 
shown in Table 3. 


Table 3. Relationship between web thickness, pressure, 
and velocity for banded (AT) and pre-engraved (PE) 
caliber .50 projectiles. 


Bullet 

Powder 

charge 

(grains) 

Web 

thickness 

(in.) 

Pressure 

(psi) 

Velocity 

(fps) 

AT 

465 IMR 

0.0338 

54,000 

3,452 

PE 

465 IMR 

0.0338 

36,100 

3,160 

PE 

465 IMR 

0.0285 

53,900 

3,520 



1 COPPER BANDED ARTILLERY TYPE 

2 PE STEEL 


Figure 4. Comparison of pressure and velocity change 
for artillery banded and PE steel projectiles, fired in the 
caliber .50 erosion-testing gun with double-base powder 
at a muzzle velocity of 3,700 fps. 


Pressure-Velocity Change and Bore 
Enlargement 

Any change in the forcing cone or other area of the 
bore close to the origin of rifling affects the starting 
resistance of an engraving-type bullet. This conse¬ 
quently affects the pressure and the velocity. How¬ 
ever, the engraving component of the starting resist¬ 
ance has been completely eliminated when using PE 
bullets. This means that the pressure and the velocity 
of a PE projectile are little affected by any enlarge¬ 
ment at the origin of rifling. 

Experiments have shown that this is not actually 
true. As the bore enlarges due to gas erosion, the 
amount of gas leaking past the bullet increases. Dur¬ 
ing the early stages of the burning of the powder, the 
rate of burning is seriously affected by any change in 
the amount of gas that might leak past the bullet and 
thus affect the pressure over the powder. For this 
reason, as the bore enlarges there is a gradual decrease 
in pressure and velocity. The use of obturating devices 
to prevent this gas leakage and maintain a more uni¬ 
form pressure and velocity level is described in a later 
paragraph of this section. 

A comparison of the pressure and velocity per¬ 
formance with artillery-type bullets having copper 
rotating bands (designated as AT) and PE bullets is 
shown in Figure 4. These data are an average of two 
tests. From these curves it may be seen that a velocity 
drop of 200 fps occurred much sooner with the en¬ 
graving-type bullets and the pressure drop corre¬ 
sponding to that velocity change was 8,800 psi 
instead of 7,200 psi. 


Effect of Double Engraving on Powder 
Pressure 

As the lands at the origin of rifling erode and the 
point of engagement with the PE projectile advances 
towards the muzzle there comes a time when the PE 
projectile does not mesh with the rifling and engraving 
of the steel band might occur. Since the powder web 
and load were adjusted for no engraving resistance 
there was a very good chance of obtaining excessively 
high pressures and producing damaging effects if 
double engraving occurred. However, examination of 
the curve of pressure versus rounds fired in Figure 4 
shows a continuous drop in pressure, even though 
double engraving of the steel band occurred between 
rounds 215 and 285. It is believed that the gas leakage 
which occurred due to the increase in bore dimensions 
caused by erosion acted as a safety valve when the 
bullet was momentarily retarded while being double 
engraved. 

Effect of Obturation on Velocity Performance 

To determine the effect of various means of obtura¬ 
tion on velocity performance of PE projectiles the 
following obturated caliber .50 PE projectiles shown 
in Figure 5 were tested: 

1. Standard steel PE projectile. 

2. Standard PE projectile with copper obturating 
ring behind rear PE band. 

3. Standard PE projectile with a cellulose acetate 
(plastic) obturating cup attached to the base of the 
projectile. 


CONFIDENTIAL 


























































596 


PRE-ENGRAVED PROJECTILE WITH CHROMIUM-PLATED BORE 


1- STANDARD PRE-ENGRAVED PROJECTILE NOT OBTURATED 

.508" DIA 



2- COPPER RING OBTURATED 


COPPER RING .508" DIA 



3- PLASTIC CUP OBTURATED 


PLASTIC CUP .508" DIA 



4* DOUBLE OBTURATED 



Figure 5. Types of obturated PE projectiles used in 
tests with the caliber .50 erosion-testing gun. 


4. Standard PE projectile combining the copper 
obturating ring with the plastic obturating cup. 

To test the effectiveness of the obturating devices 
velocity measurements were taken at various stages 
in the life of the gun. The erosion rounds during the 
test using Firing Schedule II were standard PE bul¬ 
lets (No. 1 in Figure 5). The initial firing conditions 


are given in Table 4, and the velocity changes at 
various stages of bore erosion are shown in Figure 6. 
In all instances the obturating devices increased the 
number of rounds fired for a velocity drop of 200 fps. 

These data show that the velocity performance can 
be improved and maintained at a more uniform level 
by using obturating devices with PE projectiles. The 
best obturating device was that used with projectile 
No. 4. The copper obturating ring was effective at 
low powder pressure and the plastic base cup be¬ 
came effective as the powder pressure reached its 
maximum. 



<3 



O 50 100 150 20 0 250 3 0 0 350 4 00 

ROUNDS FIRED 


1 STANDARD PE 

2 COPPER RING OBTURATED 

3 PLASTIC CUP 

4 COPPER RING AND PLASTIC CUP OBTURATED 

Figure 6. Velocity performance of the obturated PE 
projectiles shown in Figure 5, fired at a muzzle velocity 
of 3,650 fps. 


Table 4. Comparison of velocity change with obturated caliber .50 PE projectiles. 


Bullet No. 

Powder (476 grains) 

1 

30 Per cent slow 

70 Per cent fast 

2 

75 Per cent slow 

25 Per cent fast 

3 

75 Per cent slow 

25 Per cent fast 

4 

75 Percent slow 
25 Per cent fast 

Initial pressure (psi) 

56,700 

53,800 

47,700 

56,200 

Initial velocity (fps) 

3,645 

3,650 

3,440 

3,673 

Velocity after 355 rd (fps) 

3,340 

3,295 

3,180 

3,363 


CONFIDENTIAL 



























































































































CALIBER .50 EROSION-TESTING GUN EXPERIMENTS 


597 


314 4 Effect of Pre-Engraving on Accuracy 

The superior performance of the PE steel projectiles 
in a new and eroded barrel is due entirely to the high 
strength of the material in the band. Equally good 
accuracy has been obtained with a bullet having a 
steel band which has to be engraved in the gun. In 
the latter case, however, severe abrasion of the barrel 
occurred due to the high friction and engraving stresses 
of the steel band. 

A comparison of the accuracy of ball bullets, M2, 
copper-banded artillery-type bullets and PE steel 
bullets fired for a range of velocities in new and eroded 
caliber .50 barrels is shown in Table 5. These data are 


Table 5. Accuracy of different types of caliber .50 
bullets. 


Mean radius of dispersion (in in.) at 100 ft 


Velocity 

(fps) 

Ball, M2 

Banded 
artillery type 

PE 

2,900 ± 50 

0.4 

(a) In new barrel 

0.5 

0.45 

3,300 ± 50 

1.2 

1.0 

0.8 

3,700 ± 50 

7.9 

1.1 

0.7 

2,900 ± 50 

1.2 

(b) In eroded barrel 
0.75 

0.45 

3,500 ± 50 

10.9 

1.45 

0.45 


shown graphically in Figures 7 and 8. These results 
show that: (1) ball bullets, M2, are not satisfactory 
for high velocity; (2) a banded artillery-type projec- 



29 0 0 33 0 0 3700 

VELOCITY IN FPS 

Figure 7. Dispersion of ball, M2, banded AT and PE 
projectiles for various velocities fired in the caliber 
.50 erosion-testing gun using a new steel barrel. 



Figure 8. Dispersion of ball, M2, banded AT and PE 
projectiles for various velocities fired in the caliber .50 
erosion-testing gun using an eroded barrel. 

tile is intermediate in accuracy between the ball and 
PE projectile and its accuracy is affected by erosion 
of the bore; (3) a PE steel projectile is the most 
accurate of these bullet types, and its accuracy is not 
affected by erosion of the bore. 

In firing gun steel barrels with PE bullets until 
accuracy failure occurs the accuracy of the PE bullet 
remains practically unchanged until engagement of 
the teeth on the bullet with the rifling in the barrel is 
lost. When misalignment occurs, the steel PE bullet 
is double engraved, and severe wear of the lands 
occurs. During this period the accuracy remains the 
same until a bore diameter is reached when the bear¬ 
ing area in contact with the lands is greatly overloaded 
and shearing of the steel band occurs. 

This behavior results in a very uniform, good ac¬ 
curacy pattern throughout the accuracy-life of the 
gun, after which it suddenly becomes very bad. This 
is shown clearly in Table 6. The targets in Figure 9 
illustrate the progress of accuracy failure. It is readily 
seen that (1) the accuracy is uniformly maintained 
throughout the life of the gun, (2) there is no gradu¬ 
ally increased dispersion, and (3) accuracy failure is 
very bad when it does occur. 


CONFIDENTIAL 






















































598 


PRE-ENGRAVED PROJECTILE WITH CHROMIUM-PLATED BORE 


314 5 Effect of Band Material on 
Projectile Performance 

The strength of the band material and the type of 
surface riding on the bore surface are important fac¬ 
tors in the performance of a PE projectile. The strength 
of the band material (strength factor) is usually the 
determining factor in the accuracy-life of the gun, 
and the type of surface on the PE band and bourrelet 
(friction factor) is an important factor in decreasing 


the wear of the lands and grooves and thereby in¬ 
creasing the velocity-life of the gun. 

Strength Factor. A comparison of the test data, such 
as shown in Figure 10, obtained by firing steel-banded 
PE and gilding metal-banded PE projectiles, shows a 
reduction in the land and groove wear when using 
gilding metal PE projectiles. The pressure and veloc¬ 
ity drop caused by gilding metal-banded PE projec¬ 
tiles, however, are identical with those caused by the 
steel PE projectiles. 



Figure 9. Accuracy targets for caliber .50 PE projectiles at 100 feet, during the firing of 54-round groups at four stages 
in the life of a steel barrel in the erosion-testing gun. The barrel reached the end of accuracy-life during the last group. 


CONFIDENTIAL 










CALIBER .50 EROSION-TESTING GUN EXPERIMENTS 


599 



artillery-type, steel PE, and gilding metal PE pro¬ 
jectiles fired in the caliber .50 erosion-testing gun at a 
muzzle velocity of 3,700 fps. 

A comparison of the accuracy measurements of 
artillery banded, steel PE, and Parco-Lubrized steel 
PE projectiles during the erosion life of a gun steel 
barrel is given in Table 6. 


Table 6. Increase in accuracy of caliber .50 pre-en- 
graved projectiles compared with copper banded artil¬ 
lery type fired at 3,700 fps from a new barrel. 


Rounds 

Mean radius of dispersion (in in.) at 100 ft 

Artillery 

type 

Gilding 
metal PE 

Steel 

PE 

Parco- 
Lubrized 
steel PE 

1- 10 

1.0 


0.85 

1.0 

11- 65 

2.4 

1.4 

1.1 

1.0 

81-135 

3.3 

1.6 

1.1 

0.9 

151-205 


2.1 

1.1 

1.0 

221-275 


4.0 

1.1 

1.0 

291-345 



3-2* 

1.0 

370-424 




0.7 

440-494 




2.6* 


* Keyholing bullets. 


Keyholing bullets occurred after 280-290 rounds 
for the steel PE projectiles, and after 425-435 rounds 
for the Parco-Lubrized steel PE projectiles, thereby 
showing a one and one-half increase in accuracy life. 

It is clearly seen that the progressive failure of the 
gilding metal band as erosion increases is due entirely 
to the lower strength of the band material. For this 
reason gilding metal is not recommended as a band 
material when fired under hypervelocity conditions. 

Friction Factor. The decrease in land wear due to a 
change in coefficient of friction from steel against 
steel to gilding metal against steel is shown in Figure 
10. Curve 1 gives the land enlargement for the en¬ 
graving copper-banded projectiles. Curve 2 gives the 
same for the steel PE projectiles. Therefore, the area 
A between curves 1 and 2 represents the amount of land 
wear caused by the engraving of the rotating band. 


Curve 3 gives the land enlargement for the gilding 
metal PE projectiles. Therefore, the area B between 
curves 2 and 3 represents the amount of land wear 
contributed by the friction of the steel-banded PE 
projectile. 

In another pair of tests, steel PE projectiles were (1) 
cadmium plated, and (2) Parco-Lubrized c with oil 
treatment and then tested for accuracy and velocity 
performance and reduction in land and groove erosion. 

Compared with the plain steel PE projectiles less 
land wear was observed with the cadmium-plated and 
Parco-Lubrized bullets. However, best all-around 
performance was observed with the Parco-Lubrized 
and oil-treated projectiles. 

The comparison of land wear between banded ar¬ 
tillery-type, steel PE and Parco-Lubrized PE projec¬ 
tiles is shown in Figure 11 and Table 7. Area A 
between curves 2 and 3 represents the amount of land 
wear at 0.5 in. from the origin of rifling contributed 
by the friction of the steel-banded PE projectile. 



Figure 11. Comparison of land wear for banded artil¬ 
lery-type, steel PE, and Parco-Lubrized PE projectiles 
fired in the caliber .50 erosion-testing gun at a muzzle 
velocity of 3,650-3,700 fps. 

The reduction in wear of the lands and grooves is 
reflected in the increase in velocity-life observed with 
barrels fired with Parco-Lubrized steel PE projectiles. 
The observed velocity changes at various stages of 
erosion, compared with those for banded artillery- 
type projectiles, are shown graphically in Figure 12. 
Using a 200-fps drop in velocity as the criterion for 
the velocity-life of the gun, these results, which are 
summarized in Table 8, show a twofold increase in 
velocity-life by Parco-Lubrizing the steel PE projec¬ 
tiles. The same increase in velocity-life of a chromium- 
plated bore surface was also observed. 


c Defined in footnote (k) in Section 27.3.4. 


CONFIDENTIAL 

























































PRE-ENGRAVED PROJECTILE WITH CHROMIUM-PLATED BORE 


600 


Table 7. Land wear of caliber .50 barrels fired with pre-engraved projectiles at 3,650 to 3,700 fps (see Figure 11). 



Distance 
from O.R. 
(in.) 



Increase in land diameter 
(10~ 3 in.) 



Bullet 

After rounds 70 

140 

210 

280 

350 

500 

Steel PE 

Steel PE 

Parco-Lubrized PE 
Parco-Lubrized PE 

0.5 

2.0 

0.5 

2.0 

6.4 

3.1 

3.3 

2.9 

11.4 

7.8 

4.9 

3.3 

18.1 

14.4 

8.1 

6.8 

22.5 

18.8 

11.5 

10.0 

14.5 

12.5 

20.5 

18.2 



Figure 12. Comparison of velocity change for banded 
artillery-type, steel PE, Parco-Lubrized steel PE pro¬ 
jectiles fired in the caliber .50 erosion-testing gun at a 
muzzle velocity of 3,650-3,700 fps. 


obtained when they were fired with those when banded 
artillery-type projectiles were fired, it was possible to 
distinguish between the amount of wear of the lands 
contributed by engraving of the band and friction of 
the band. As shown in Figure 9 of Chapter 15, dis¬ 
cussed in Section 15.5.2, this amount is propor¬ 
tionately much higher for propellants having low 
flame temperatures, for the powder gas erosion be¬ 
comes the predominating factor with the hotter pro¬ 
pellants. 

3i 5 OPTIMUM THICKNESS 

OF CHROMIUM PLATE 


Table 8. Velocity life of caliber .50 steel barrels fired 
with bullets having rotating bands of different materials. 


Initial Velocity- 

velocity life* 

Bullet (fps) (rounds) (relative) 

Artillery-type (copper banded) 3,657 95 1.0 

Gilding metal banded, PE 3,712 215 2.3 

Steel, PE 3,703 225 2.4 

Steel, PE, Parco-Lubrized 3,646 455 4.8 

* Number of rounds at which velocity drop equaled 200 fps. 

314 6 Effect of Type of Powder 

In the determination of the erosiveness of propel¬ 
lants, which is described in Chapter 15, PE projectiles 
played an important role. By comparing the results 


Altered layers, similar to those formed in unpro¬ 
tected steel (Section 12.1.2), are observed in the steel 
underlying thin chromium plates. Since the formation 
of the altered layer is a thermal effect, the thickness 
of the altered layer produced is a function of the plate 
thickness and of the temperature and heat content 
of the powder gases. Experimental data, obtained 
with caliber .50 and 37-mm guns, suggest that there 
is a definite chromium plate thickness for each gun 
that will give the maximum performance of PE pro¬ 
jectiles. This critical chromium plate thickness is 
determined by the heat input to the bore surface. In 
other words, guns having a high heat input per unit 
area require a thicker chromium plate than guns 
having a low heat input. 



Figure 13. Effect of powder type and chromium-plate thickness on performance of PE projectiles fired in the caliber 
.50 erosion-testing gun. 


CONFIDENTIAL 


























































































INCREASE IN LIFE OF CHROMIUM-PLATED BORES 


601 


In order to vary the heat input per unit area in the 
caliber .50 erosion-testing gun, powders of different 
flame temperature were used. The thickness of the 
chromium plate was varied from 0.002 to 0.006 in. In 
all tests the same ballistic level was maintained; 
namely, a pressure of 56,000 to 58,000 psi (copper) 
and a velocity of 3,600 to 3,650 fps. The powders used 
were (1) IMR type, flame temperature 2940 K; (2) 
FNH-M2, flame temperature 3560 K; (3) a double¬ 
base powder containing 40% nitroglycerin, flame 
temperature 3945 K. 

Standard steel-banded PE projectiles (Figure 1) 
were used in the tests. The barrels were fired until a 
drop in velocity of 200 fps was observed. The velocity- 
life of the barrels having the various chromium plate 
thicknesses and fired with the different powders is 
summarized in Table 9. 


Table 9. Velocity-life of caliber .50 barrels firing PE 
projectiles as a function of chromium plate thickness 
and powder. 


Thickness of Cr plate 

Rounds fired for a velocity 

(in.) 

drop of 200 fps 

(a) Powder: 40% (nitroglycerin content) double-base; 

heat input, 19.6 cal/cm 2 . 

0 

90 

0.003 

120 

0.004 

130 

0.00625 

180 

(b) Powder: FNH-M2 (double-base containing 20% 

nitroglycerin); heat input, 17.1 cal/cm 2 . 

0 

146 

0.0035 

470 

0.0045 

490 

0.005 

765 

(c) Powder: IMR (single-base); heat input, 14.4 cal/cm 2 . 

0 

362 

0.00225 

710 

0.004 

1,165 

0.006 

2,970 


The above data are shown graphically in Figure 13. 
These curves show that when the heat input is above 
19 cal/cm 2 , a chromium plate thickness greater than 
0.006 in. is necessary in order to obtain the best 
velocity performance of PE projectiles. When the heat 
input is between 14 and 15 cal/cm 2 the best velocity 
performance is observed with 0.006-in. chromium 
plate thickness. When the heat input is about 17 cal/ 
cm 2 , good performance is observed with 0.005-in. 
chromium plate, but considerably better perform¬ 
ance would have been observed if the plate thickness 
were between 0.007 in. and 0.008 in. 


It has been shown 76 in the testing of various chro¬ 
mium plate thicknesses that there is a limit to the 
thickness that will perform satisfactorily due to the 
brittle nature of chromium. Above 0.008-in. thickness, 
chromium plate fired with PE bullets breaks away on 
the lands and exposes gun steel or leaves a thin layer 
of plate, both of which are very susceptible to gas 
erosion. 

A thick cobalt undercoat (0.007 to 0.010 in.) next 
to the steel followed by a thinner chromium layer 
(0.003 in.) performed very satisfactorily in some pre¬ 
liminary firing tests under hypervelocity conditions, 
using PE projectiles, as described in Section 20.2.4. 

It is recommended, therefore, that when calcula¬ 
tions of heat input in a particular gun are very high 
and show the need of a chromium plate thicker than 
0.008 in., that a thick cobalt-chromium duplex plate 
be tried. 

316 INCREASE IN LIFE OF CHROMIUM- 
PLATED BORES 

In the early firings of chromium-plated bores with 
the usual engraving-type projectiles it was soon rec¬ 
ognized that engraving stresses and friction were a 
major factor in the failure of the chromium plate. 
Detailed description of the progress of chromium 
plate failure under hypervelocity conditions is given 
in the final report on coatings. 76 The process of erosion 
in chromium-plated Service guns is described in Sec¬ 
tion 20.2.1. 

In general, the failure of a chromium-plated surface 
is due to (1) engraving stresses, (2) bullet friction, 
and (3) thermal changes at the gun steel-chromium 
interface. The first two items are effective chiefly 
because of the lack of ductility of the chromium. 

The features of chromium plate failure usually 
observed are: (1) cracking of the plate in a block or 
checkerwork pattern; (2) curling up of the edges of 
the blocks which gives the surface a wrinkled appear¬ 
ance ; (3) pitting, or removal of small crack-isolated 
blocks of chromium; (4) spalling or removal of large 
areas due to undercutting; and (5), in the case of 
plates thinner than the critical thickness, the forma¬ 
tion of an altered steel layer beneath the plate. 

As mentioned in Section 31.4.4, the wear and fric¬ 
tion factors in the erosion of gun steel were greatly 
reduced by the use of PE projectiles. Since chromium 
plate, because of its high melting point (about 1950 C), 
possesses excellent resistance to powder gas erosion, 
the combination of a chromium-plated bore and PE 


CONFIDENTIAL 







602 


PRE-ENGRAVED PROJECTILE WITH CHROMIUM-PLATED BORE 



Figure 14. Progress of land erosion of plain steel and chromium-plated steel barrels in the caliber .50 erosion-testing gun 
fired with banded artillery-type, steel PE, and Parco-Lubrized PE projectiles at a muzzle velocity of 3,675 to 3,700 fps. 


Parco-Lubrized steel projectiles gives excellent per¬ 
formance. 

3161 Effect on Land Erosion 


erosive effects of the powder gases, the rate of erosion 
increases very rapidly. Parco-Lubrized PE bullets 
give the longest protection period and the lowest 
erosion rate after the gun steel has been exposed. 


In the following tests double-base powder was used 
to give a velocity of 3,650 to 3,700 fps in a 45-in. 
barrel. The observed increase in land diameter at 0.5 
in. beyond the origin of rifling for gun steel and 0.005- 
in. chromium plate using banded artillery-type and 
steel PE projectiles is shown graphically in Figure 14. 
The results show that the chromium plate protects 
the steel surface for a definite period. This period of 
protection is much shorter for the banded artillery- 
type bullets. As soon as gun steel is exposed to the 


316 2 Effect on Velocity Performance 

The velocity change observed with banded and PE 
projectiles fired in gun steel and chromium-plated 
bores is summarized in Figure 15. 

Using a drop in velocity of 200-fps as the criterion 
for the velocity-life of the gun, it is readily seen that 
Parco-Lubrized PE bullets combined with a chro¬ 
mium-plated bore improves the velocity performance 
of the gun. The performance is shown in Table 10. 



Figure 15. Comparison of velocity change for banded artillery-type, steel PE and Parco-Lubrized PE projectiles 
fired from plain steel and chromium-plated steel barrels in the caliber .50 erosion-testing gun at a muzzle velocity of 
3,675-3,700 fps. 


CONFIDENTIAL 











































































































DESIGN AND TEST OF THE 37-MM GUN, T47 603 



Table 10. Comparison 

of velocity performance. 




* 

Initial 

. 

Rounds 

Relative 



velocity 

fired for 

velocity- 

Bore surface 

Bullet 

(fps) 

AV = —200 fps 

life 

Gun steel 

Artillery banded 

3,675 

95 

1.0 

Gun steel 

PE 

3,703 

225 

2.4 

0.005-in. Cr plate 

Artillery banded 

3,700 

220 

2.3 

0.005-in. Cr plate 

PE 

3,700 

975 

10.0 

0.005-in. Cr plate 

Parco-Lubrized PE 

3,675 

1,875 

20.0 


The decrease in wear and friction of the Parco- 
Lubrized bullets on a chromium-plated bore is shown 
in the increase in velocity-life of the gun. These results 
show that in order to obtain the maximum perform¬ 
ance with PE projectiles it is necessary to (1) Parco- 
Lubrize all parts of the projectile that rub on the 
bore, and (2) use a chromium-plated bore having a 
chromium-plate thickness greater than the critical 
thickness necessary to prevent formation of an altered 
steel layer at the steel-chromium interface. 

316 3 Effect on Accuracy Performance 

In one test, Parco-Lubrized PE bullets were fired 
in a chromium-plated bore (0.005-in. chromium plate) 
until key holing bullets were observed at 100 ft from 
the muzzle. This did not happen until 2,775 rounds 
had been fired, whereas with a gun-steel bore and 
steel PE bullets, keyholing started after 248-322 
rounds had been fired. Thus the combination of chro¬ 
mium plate and Parco-Lubrized steel PE bullets in¬ 
creased the accuracy-life by 8.5 to 10 times. 


317 DESIGN AND TEST OF THE 37-MM 
GUN, T47 

31,71 General Plan 

It has been shown in the preceeding sections that 
a large increase in velocity-life could be obtained in a 
hypervelocity gun by using Parco-Lubrized steel PE 
projectiles in combination with a chromium-plated 
bore. The thickness of the chromium plate should be 
such that the temperature at the chromium-gun steel 
interface is below the transition temperature of the 
steel in order to prevent the formation of an altered 
layer. 

In the caliber .50 erosion-testing gun, twentyfold 
increase in velocity-life was observed. Although the 
increase in velocity-life would not necessarily be of 


the same order for larger caliber guns, it was felt that 
it would be substantial. 

As a preliminary to the design of a hypervelocity 
90-mm gun (Section 31.8), it was decided to make a 
hypervelocity 37-mm gun and observe the perform¬ 
ance of the PE chromium plate combination in this 
gun. With this end in view the following specifications 
for the gun tube were set up. 

1. A 40-mm Bofors gun forging (Navy specification) 
length: 88.58 in. 

2. Chambered for 40-mm case necked down to 
37-mm. (Chamber capacity 29.9 cu in.) 

3. Bored and reamed for a 37-mm projectile. 

4. Rifled for 6 grooves of 0.0030-in. depth, with lands 
and grooves of equal width. 

5. Lands at the origin of rifling pointed with a 
30-degree slope to eliminate “hang-ups.” 

6. Plated with 0.006-in. standard chromium. 

By firing a normal weight pre-engraved projectile 
at a pressure of 50,000 psi (copper), it was expected 
that a muzzle velocity of 3,500 fps would be achieved. 

317,2 Gun Tube 

The 40-mm gun forging was chambered, bored, 
and rifled according to the dimensions shown in 
Figure 16. 

The constants for the 37-mm gun, T47, are as 


follows: 


Length 

88.58 in. 

Chamber capacity 

29.9 in. 3 

Travel of projectile 

76.58 in. 

Number of grooves 

6 

Depth of grooves 

0.030 in. 

Twist 

1 in 25 calibers 

Ratio: width of groove/shelling 


width of land 

1 

Land diameter (after plating) 

1.457+.002 in. 

Groove diameter (after plating) 

1.517 + .002 in. 

Groove width (after plating) 

0.375+.005 in. 


CONFIDENTIAL 







604 


PRE-ENGRAVED PROJECTILE WITH CHROMIUM-PLATED BORE 


. 375" CHORD 
AFTER PLATING 
(MACHINE TO.380" 



Figure 16. Chamber and rifling of a 40-mm forging used for the 37-mm gun, T47, to fire PE projectiles. 


Stress calculations were made for the chamber, 
which was the critical section of the tube. The results 
showed that the pressures to be used in firing were 
dangerously near the yield point of the steel. How¬ 
ever, since PE projectiles were being used, the time 
of maximum pressure was reduced and the radial load 
due to engraving was also eliminated; therefore it 
was felt that the gun tube was strong enough to with¬ 
stand 50,000 psi (copper). Two gun tubes were later 
proof-fired at 62,000 psi (copper) with no serious 
trouble. 

Chromium plating of the bore surface was to start 
at a point 11% in. from the breech face and proceed 
to the muzzle. Allowance in the land dimensions was 
made so that after machining, 0.002 in. on the radius 
was to be removed by electropolishing, to give a final 
land diameter before plating of 1.469 (+.002) in. The 
application of 0.006-in. chromium plate would then 
give the specified land diameter for the finished tube. 
Allowance in the groove dimensions was made for a 
25 per cent less distribution of the plate thickness 
than was to be deposited on the lands. 

The groove dimension was made smaller because 
it was expected that less chromium would be de¬ 
posited on the groove surface than on the land surface. 
This difference in plate distribution was believed due 
to the difference in distances from the anode. Previous 
experience indicated that when 0.006-in. chromium 
was deposited on the land surface, about 0.0045-in. 


chromium was deposited on the groove surface. The 
application of 0.0045-in. chromium plate on the 
grooves would give a finished groove diameter of 
1.517 in. ( + .002 in.) 

After plating, the gun tubes were found to be con¬ 
stricted in the grooves. The grooves were lapped to 
1.516-in. diameter so that a PE projectile, whose 
maximum diameter across the teeth was 1.513 in., 
easily passed through the bore. The final plate thick¬ 
ness was found by gauging to be only 0.0045 to 0.0055 
in. on the lands and 0.0030 to 0.0035 in. on the 
grooves. 

3173 Ammunition 

Special ammunition components were also required. 
A PE projectile weighing 1.62 lb and made accord¬ 
ing to Figure 17 was used as the standard projectile 
to obtain a velocity of 3,500 fps. Other projectiles 
weighing 1.34 lb and 1.92 lb were also made. The 
standard projectile and assembled round are shown 
in Figure 18. 

The standard 40-mm cartridge case, M22A1, was 
resized at the neck and shoulder so that a 37-mm PE 
projectile would fit. 

Firings to establish the powder granulation and 
load were made at the Ordnance Research Center, 
Aberdeen Proving Ground. 240 The results of these 
firings are given in Table 11. 


CONFIDENTIAL 












































605 


DESIGN AND TEST OF THE 37-MM GUN, T47 


Table 11 . Powder granulation and load for 37-mm gun, T47. 240 


Proj. 

wt. 

(lb) 


Powder 


Muzzle velocity (fps) 

Pressure (psi) 

Lot 

Type 

Web 

(in.) 

Charge 

(oz) 

Mean 

Max. 

var. 

Mean 

dev. 

Mean 

Max. 

var. 

Mean 

dev. 

1.34 

13231 

M5 

0.0400 

12.10 

3,486 

22 

5.9 

35,100 

3,300 

700 

1.34 

13231 

M5 

0.0400 

12.82 

3,668 

24 

8.7 

40,900 

2,200 

600 

1.62 

13225 

M5 

0.0400 

13.10 

3,493 

17 

4.0 

43,300 

3,900 

1,000 

1.62 

13231 

M5 

0.0400 

12.82 

3,495 

62 

13.7 

44,200 

3,300 

900 

1.62* 

13231 

M5 

0.0400 

12.54 

3,451 

34 

6.7 

43,900 

2,600 

700 

1.62* 

13231 

M5 

0.0400 

12.82 

3,525 

20 

5.3 

47,700 

5,400 

1,200 

1.62 

14822 

Ml 

0.0231 

13.25 

3,378 

45 

11.9 

49,800 

2,900 

900 

1.92 

Pilot 299 

M2 

0.0425 

12.00 

3,284 



49,800 

' 


1.92 

13231 

M5 

0.0400 

12.82 

3,344 

28 

7.9 

52,400 

2,800 

700 


* Projectile with obturating disk. 


The results of these tests showed that a powder 
having the M5 composition and a web of approxi¬ 
mately 0.0400 in. was suitable as a propellant to yield 
a muzzle velocity of 3,500 fps within a pressure limit 
of 50,000 psi (copper) when using 1.34- and 1.62-lb 
projectiles, with or without an obturating disk. 

ASSEMBLED 


317 4 Erosion Test 

One of the two 37-mm gun tubes, T47, was mounted 
on a modified 57-mm carriage and sent to Aberdeen 
Proving Ground for an erosion test. The other one, 
similarly mounted, was sent to Division 1, NDRC 

PROJECTILE 


SHARP CORNERS 



Figure 17. PE projectile for 37-mm gun, T47: Solid steel body with aluminum windshield, unobturated, weight 1.62 lb. 


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606 


PRE-ENGRAVED PROJECTILE WITH CHROMIUM-PLATED BORE 



Figure 18. 37-mm PE projectile and assembled 
round. 


ballistic firing range at Carderock, Maryland, for the 
ballistic tests summarized in Chapter 4. 

For the erosion test, a charge of 12.85 oz of FNH-M5 
powder was established to give a velocity of 3,500 fps 
with a pressure of 43,500 psi (copper) using 1.62-lb 
PE projectiles. Erosion rounds were fired in groups of 
50 rounds at a rate of fire of 5 to 6 rounds per minute. 

Failure of the plate occurred in the grooves and the 
subsequent stripping of the plate and erosion of the 
exposed gun steel caused a drop in velocity of 200 fps 
after 541 rounds. This failure of the chromium plate 
can be attributed to too thin a plate (less than 0.006 
in. on both lands and grooves—see Section 31.7.2) to 
prevent alteration of the steel at the chromium plate- 
steel interface. It has been shown in Section 31.5 that 
chromium plate thickness is the determining factor 
in obtaining long gun life, when PE projectiles are 
fired from a chromium-plated bore. 

It is, therefore, recommended that another 37-mm 
gun tube T47 be plated with the proper thickness 
(0.006 in.) of chromium plate on the lands and grooves 
to insure thermal protection of the steel at the inter¬ 
face. 

318 DESIGN OF CHROMIUM-PLATED 

HYPERVELOCITY 90-MM GUN TO FIRE 
PRE-ENGRAVED PROJECTILES d 

General Plan 

By the time that Division 1 had made considerable 
progress in the development of erosion-resistant ma¬ 
terials, it was realized that their eventual application 
to hypervelocity guns of medium caliber would re¬ 
quire testing them in an experimental hypervelocity 
gun. Such a gun would also give an opportunity to 
explore directly some of the ballistic factors believed 
to be important in connection with the design of a 
hypervelocity gun—especially by the use of very 
high densities of loading. 

The project for the development of this gun was 
termed the A-Z project, because it was expected that 
all phases of the Division’s program (“from A to Z”) 
would be drawn upon as it grew. Thus in the end the 
series of A-Z guns would be the embodiment of the 
Division’s accomplishments. The material selected 
for the first trial was chromium electroplate, to be 
used in conjunction with pre-engraved projectiles. It 


d This section has been added by the Editor to round out the 
subject of this chapter. It is based on an NDRC report. 118 


31.8.1 


CONFIDENTIAL 







CHROMIUM-PLATED HYPERVELOCITY 90-MM GUN 


607 


was understood that as other materials became avail¬ 
able, other tubes would be prepared to test them in a 
gun of this size, and of the same general design. 

As a guide to the choice of specifications for this 
gun, an extensive series of ballistic calculations was 
made with the Division 1 system; the results are 
described in Section 3.5.5. The choice was finally 
narrowed down to a gun having a muzzle velocity 
slightly in excess of 4,000 fps and an overall length of 
not more than 90 calibers. 

31,8-2 Gun Tube Design 

Further ballistic calculations led to adoption of the 
specifications listed in Table 12. From them a strength 


Table 12. Basic specifications for 90-mm gun, A-Z. 


Bore diameter (90-mm) 

3.524 in. 

Length of gun 

88 calibers 

Weight of projectile 

23.8 lb 

Muzzle velocity 

4,200 fps 

Maximum powder pressure 

60,000 psi 

Volume of chamber 

760 cu in. 

Density of loading 

0.80 

Powder type (7-perforate) 

FNH-M1 

Travel of projectile 

281.0 in. 


curve was calculated in thepsualway (Section 26.4.4), 
with the results shown in Figure 19. 

One requirement determining the choice of speci¬ 
fications for the tube was that in the experimental 
firing the 120-mm gun mount, Ml be used in order to 
save the time required to make a special mount. Also, 
consideration was given to the possibility that the 
120-mm breech ring and breech block be used with 
minor modifications. 

The design of the chamber involved several com¬ 
promises. Its total volume was fixed by the basic 
specifications. A long slender case was preferable as 
far as strength of the tube was concerned, and, in fact, 
the requirement that the 120-mm breech ring be used 
imposed an upper limit on the diameter of the cham¬ 
ber at the rear. As far as ease of fabrication of the 
cartridge case and ease of manipulation by an auto¬ 
matic loading mechanism were concerned, a short 
case was favored. The compromise made between 
these two competing requirements led to a chamber 
that was 34 in. long and 6.0 in. in diameter at the 
rear. This diameter corresponds to a wall ratio of 2.6. 

It was decided to make the cartridge case of brass, 
because of uncertainties regarding the likelihood of 
steel cases being made with the high strength neces- 



Figure 19. Strength curve and pressure-travel curve for tube of 90-mm gun, A-Z. 


TLnMJ/JJi _ 



znzdLlll --i- 




CONFIDENTIAL 




















































608 


PRE-ENGRAVED PROJECTILE WITH CHROMIUM-PLATED BORE 


sary to avoid sticking. Special attention was given in 
the design to the avoidance of cracking close to the 
inside corner of the base. 

The forging design adopted was a modification of 
the forging for the 120-mm gun tube. The finished 
exterior dimensions at the rear were made identical 
with those for the 120-mm tube, so that the tube 
would fit the mount. The entire tube was somewhat 
longer (315.0 in. instead of 283.5 in.) than the 120- 
mm, and the forward end was made with a smaller 
external diameter. With this design the maximum 
strain in the tube at the origin of rifling was 0.0028 
in. per in. and it was 0.0031 in. per in. at the large 
chamber cone. These values of strain are close to that 
which separates a critically stressed tube from a 
normally stressed one. 

318 3 Gun Tubes 

The A-Z project was halted by the termination of 
Division Ts activities. The preparation of the tube 
forgings, however, was far enough along so that it 
seemed expedient to finish them. Six forging blanks 
had been cast from an alloy steel containing nearly 
3% nickel, about 0.9% chromium, about 0.5% 


molybdenum, and about 0.1% vanadium. The mini¬ 
mum yield strength specified was 130,000 psi. 

During machining of the heat-treated forgings, 
cracks in the steel were discovered in four of them, 
and they had to be rejected. In one case the cracks 
were not disclosed until the final honing operation. A 
study of the test data, including the examination of 
microstructure, suggested that the reason for this 
high percentage of rejections was poor melting prac¬ 
tice, which resulted in steel with too high an inclu¬ 
sion rating to give the requisite degree of finished- 
surface continuity. 

The machining of two tubes was completed as far 
as the exterior, the chamber, and a smooth bore were 
concerned. These tubes, ready for rifling, were sent 
to Watervliet Arsenal in April 1946, after an agree¬ 
ment had been reached with the Army Ordnance 
Department that it would complete their manufac¬ 
ture and test them. The Naval Gun Factory had 
agreed to chromium-plate them with the cooperation 
of the Electrochemistry Section of the National 
Bureau of Standards. The original design of these 
tubes had called for pointed lands; but in view of the 
results with the 37-mm gun, T47, described in Section 
31.7.4, this decision needs to be reconsidered. 


CONFIDENTIAL 




Chapter 32 

THE FISA PROTECTOR AND CHROMIUM-PLATED BORE a 

By Nicol H. Smith b 


321 INTRODUCTION 

among the means of preventing erosion of a steel 
gun barrel, one of the simplest in principle is to 
cover the area of the bore surface just ahead of the 
origin of rifling in order to protect it from the action 
of the hot powder gases. The Fisa protector 0 is a de¬ 
vice for accomplishing this purpose. Its development 
followed information 300 received from abroad in 1942 
that plans for a new Bofors 57-mm AA gun included 
the use of a thin metal sleeve on the projectile as a 
means of controlling erosion. 

The Fisa protector is a very thin, slightly tapered 
sleeve of soft steel that is slipped over a complete 
round of ammunition, the larger end of the sleeve 
snugly fitting the neck of the cartridge case up to the 
shoulder, and the remainder covering the projectile 
almost to the bourrelet. Figure 1 shows a Fisa sleeve 
before and after firing. During firing the sleeve is 
locked to the cartridge case by the expansion of the 
neck of the case, which is extruded by the powder 
pressure into several rectangular holes cut through 
the rearward end of the sleeve, thus insuring the with¬ 
drawal of the sleeve with the case. Thus the Fisa pro¬ 
tector in effect is a replaceable steel liner (Section 
26.3) that is replaced after each round fired. 

The gun from which ammunition equipped with a 
Fisa protector is to be fired must be modified slightly 
in order to allow the round to be chambered. The 
diameters of the centering cylinder and of the bore as 
far forward as the front of the sleeve must be enlarged 
by almost twice the thickness of the sleeve. Grooves 
of standard depth are then cut in the enlarged bore to 
re-establish the rifling. 

Preliminary tests showed that the Fisa protector 
was not effective in increasing the life of a plain steel 
hypervelocity caliber .50 gun barrel, because of ex- 

a This chapter is based on a formal Division 1 report 121 by 
the same author, to which reference is made for further details. 

b The Franklin Institute, Philadelphia, Pennsylvania. 

c The term Fisa is a coined word made up of the initial letters 
of the phrase “Franklin Institute, Section A.” At the time that 
The Franklin Institute was asked to develop this device as part 
of the work under Contract OEMsr-533, that contract was 
being supervised by Section A of Division A, NDRC (later 
Division 1). 


cessive erosion of the unprotected part of the bore 
ahead of the sleeve. Further development dealt with 
the combination of a Fisa protector and a chromium- 
plated bore surface. The sleeve protects the chro¬ 
mium plate from some of the engraving stresses that 
are so detrimental to it (see Section 20.2.1), and hence 
lengthens the life of the gun barrel. In this respect, 



Figure 1 . Components, assembled round, and fired 
round for caliber .60 Fisa protector. (This figure has 
appeared as Figure 29 in NDRC Report No. A-449.) 


its function is similar to that of a pre-engraved pro¬ 
jectile, described in Chapter 31. Although the latter 
is more effective in eliminating engraving stresses, 
the former has the advantage that the round of am¬ 
munition does not need to be indexed before being 
chambered. 

Tests in the caliber .50 erosion-testing gun, which 
are described in Section 32.3, showed that the Fisa 
protector nearly doubled the velocity-life. Tests in a 


CONFIDENTIAL 


609 




610 


THE FISA PROTECTOR AND CHROMIUM-PLATED BORE 


37-mm gun, M3, revealed that the Fisa protector can¬ 
not be used with the present standard ammunition for 
that gun because of excessive interference between 
the rotating band and the grooves in the gun. These 
tests are described in Section 32.5 and some prelimi¬ 
nary ones with a caliber .60 barrel in Section 32.4. 

M.2 DESIGN OF THE FISA PROTECTOR 
32 2 1 Analysis of the Functions of the Fisa 

The most difficult aspect of the development of the 
Fisa protector was the proportioning of the different 
parts. The final design for the caliber .60 size, which 
is shown in Figure 2, includes those tapers and shoul¬ 
ders that were found to reduce to a minimum the 
various types of failure that were observed during 
preliminary firing tests. The design may be compre¬ 
hended most easily by analyzing the function of each 
of the four sections of the sleeve, which are designated 
in Figure 3. 



Figure 2. Fisa protector for caliber .60 barrel, dimen¬ 
sioned drawing. (This figure has appeared as Figure 30 
in NDRC Report No. A-449.) 


Section G 

This part of the sleeve fits over the case neck. It 
is tapered to facilitate extraction from the chamber 
after firing. Four rectangular holes are punched in 
this section to lock the sleeve to the case during fir¬ 
ing. The extrusion of the brass into these holes by the 
powder pressure performs the locking operation. 
Tapering the external dimension of this section also 
allows for a thicker section in the sleeve and thereby 
aids in a more positive locking action. 

In order to prevent failure to extract, it was found 
necessary to crimp the sleeve to the case, using the 


same crimp that is used on the projectile. This gives 
positive action at low powder pressures when the 
extrusion of the brass does not occur fast enough to 
lock the sleeve to the case in a satisfactory manner. 
Another feature that was found essential to prevent 
failure to extract was to have a carefully controlled 
clearance between section G and the centering cylin¬ 
der of the barrel. 

Section F 

The purpose of this shoulder is to prevent the sleeve 
from moving forward during the engraving operation. 
This short section is also tapered to facilitate extrac¬ 
tion of the fired cartridge case and sleeve assembly. 

Section E 

Sections E and D are the parts of the sleeve that 
are engraved. The passage of the projectile through 
this zone forces the metal in the sleeve into the rifling 
and the rotating band is thus engraved through the 
sleeve. 


ROTATING BAND FISA PROTECTOR 



Figure 3. The Fisa protector, schematic drawing. 

(This figure has appeared as Figure 5 in NDRC Report 

No. A-449.) 

The external dimension of section E is tapered rear¬ 
ward to give a thicker section having greater strength 
and also to facilitate extraction of the case and sleeve 
assembly from the chamber. The internal dimension 
of section E is also tapered in order to distribute the 
engraving stresses over a greater area. 

Section D 

This section is straight and serves as a die to finish 
the engraving operation. The purpose of this section 
was to engrave the rotating band completely so that 
the projectile would emerge from the sleeve as a fully 
pre-engraved projectile having 0.001- to 0.002-in. 
clearance between the engraved band and the rifling 


CONFIDENTIAL 









































DESIGN OF THE FISA PROTECTOR 


611 



Figure 4. Profile of caliber .50 barrel in region of origin of rifling, modified to accommodate Fisa protector. (This figure 
has appeared as Figure 6 in NDRC Report No. A-449.) 


beyond the sleeve. With regular rifling and ammuni¬ 
tion it is not possible to attain this ideal condition. 

Tearing of the Sleeve 

Tearing of the engraving parts of the sleeve (sec¬ 
tions D and E) was the most troublesome type of 
failure encountered. It was found that it could be 
minimized by: 

1. Increasing the ultimate strength of the metal in 
the sleeve. (This remedy makes fabrication by the 
deep drawing process almost impossible.) 

2. Increasing the thickness of the metal in sections 
D and E. 

3. Reducing the forward component of force from 
the moving projectile to a value below the yield point 
of the metal in sections D and E by Parco-Lubrizing 
or cadmium plating the inner surface of the sleeve. 

4. Reducing the radial load on the sleeve during 
engraving by eliminating the interference between the 
rotating band and the grooves and providing addition¬ 
al and deeper cannelures in the band, in accordance 
with the experiments described in Section 27.4.2. 

Extent of Engraving 

To determine the amount of engraving done 
through the sleeve, a caliber .50 barrel was rifled for 
a distance equal to the length of the sleeve and made 
smooth bore beyond the sleeve to the muzzle. The 
thickness of the metal in a series of sleeves and the 
inside diameters were varied. The recovered bullets 
showed that to obtain maximum engraving of the 
band it was necessary to: 


1. Keep the thickness of the metal at 0.010 in. in 
order for the engraving to be sharp, and 

2. Reduce the inside diameter of the sleeve to the 
diameter of the body of the projectile so that it would 
have to be forced over the bourrelet. 

Even by making the inside diameter of the sleeve 
for a caliber .50 bullet 0.003 in. smaller than the 
bourrelet diameter and the thickness of the sleeve 
0.010 in. it was possible to obtain only 75 to 80 per 
cent of complete engraving. This meant that addi¬ 
tional engraving of the band would occur beyond 
the sleeve, thereby subjecting the chromium-plated 
lands in that area to slight engraving stresses and 
friction. d 

32 2 3 Modification of the Rifling 

In order that the assembled round with the sleeve 
can be chambered properly, it is necessary to remove 
metal from the chamber and origin of rifling area to 
allow for the thickness of the sleeve. A profile of the 
origin of rifling modified for the caliber .50 sleeve is 
shown in Figure 4. The various sections shown match 
the sections on the sleeve (Figure 3). 

d The experience with chromium-plated caliber .50 aircraft 
machine gun barrels, described in Chapter 23, demonstrated 
that for accuracy in firing long bursts with that gun it is essen¬ 
tial to have the muzzle of the barrel slightly constricted. 
Hence it would seem undesirable to have a projectile for a 
machine gun, such as the caliber .60, completely engraved by 
the Fisa protector. On the other hand, the greater depth of 
rifling necessary with the Fisa protector might make continu¬ 
ation of engraving all the way down the bore unnecessary. In 
this connection it should be noted that in the only test of a 
Fisa protector in automatic fire (Section 32.3.4) the caliber .50 
barrel had tapered chromium plate. (Editor’s note.) 


CONFIDENTIAL 














612 


THE FISA PROTECTOR AND CHROMIUM-PLATED BORE 


32 2 4 Manufacture of the Fisa 

The first sleeves that were tested were machined 
from soft steel bar stock. Examination of the finished 
sleeves under a low-power microscope showed shallow 
grooves produced by the tools. When the thickness ot 
the sleeve was reduced to 0.010 in., these grooves pro¬ 
duced areas of weakness which resulted in tearing of 
the sleeves. 

A more satisfactory method of manufacture was 
then used to produce a better quality sleeve in greater 
quantity. The sleeves were drawn from a copper- 
coated, soft drawing steel (SAE 1010) having the 
proper thickness to give the required metal thickness 
in the finished sleeve. The copper coating prevented 
galling in the dies during the drawing operation. The 
method of manufacture was the same for the caliber 
.50, caliber .60, and 37-mm Fisas, although the num¬ 
ber of steps in the fabrication of the sleeves was dif¬ 
ferent. Figure 5 shows the steps in the fabrication of a 
37-mm Fisa protector and may be taken as repre¬ 


sentative of the manufacture of the sleeves in 
general. 

32.3 CALIBER .50 EXPERIMENTS 

32 3 1 Methods of Testing and 

Measurement 

The caliber .50 erosion-testing gun (Section 11.2.1) 
was used as the tool in the preliminary tests of the 
Fisa protector, at a velocity of 3,500 fps with a slow 
rate, of fire. The behavior of the sleeves was first 
tested in a gun steel bore and later in chromium- 
plated bores. Subsequently, they were tested in the 
Browning machine gun at a velocity of 2,900 fps and 
a cyclic rate of fire of about 600 rpm. 

All the tests in the erosion-testing gun were fired 
with double-base powder containing 20% nitro¬ 
glycerin and copper-banded, artillery-type bullets 
(Figure 3 of Chapter 11). The charge was adjusted 
to give a maximum powder pressure of 50,000 to 
52,000 psi (copper) for a velocity of 3,500 fps. 



| 2 3 4 5 <* 7 a *f XO \\ 12 



' i ' i ' i * i 1 i 1 i T i * i f i 1 i 1 » wi 1 ; 

I 2 3 4 5 <& T 8 ^ 10 11 \Z \ 

Figure 5. Operations in the fabrication of a drawn 37-mm Fisa protector. (This figure has appeared as Figure 34 in 
XDRC Report No. A-449.) 


CONFIDENTIAL 





CALIBER .50 EXPERIMENTS 


613 


The gun was considered to have failed when (1) 
the velocity had dropped 200 fps (end of velocity- 
life), or (2) the mean radius of dispersion at 100 ft had 
increased to three times its initial value or keyholing 
bullets were observed (end of accuracy-life). 

Plug gauge and star gauge measurements were 
taken at intervals to determine the distribution and 
progress of erosion of the lands and grooves. 

32 3 2 Gun Steel Bores 

The results of the tests with gun steel bores showed 
that the sleeve protected the origin of rifling against 
erosion by the powder gases, but that the erosion be¬ 
yond the origin of rifling was greater than would be 
observed normally without the use of the sleeve. This 
increased erosion beyond the sleeve is probably 
caused by increased turbulence of the powder gases 
produced by the discontinuity of the bore surface. 

Although the decrease in powder pressure was 
slight, because of the protection of the origin of ri¬ 
fling by the sleeve, there was a considerable increase 
in dispersion due to the severely eroded lands beyond 
the sleeve. This resulted in an accuracy failure of the 
gun after 55 rounds. Without the sleeve the control 
tests gave an accuracy failure of 125 rounds. 122 

32,3-3 Chromium-Plated Bores 

When the bore beyond the sleeve was protected by 
chromium plate, the accuracy failure mentioned in 
the preceding section was postponed. The gun barrel 
was plated with 0.005-to 0.006-in. standard chromium 
plate (Section 20.2.2) after electropolishing to remove 
this thickness of steel. The plate started at the 
shoulder at the breech end of section G (see Figure 
3) and was continuous to the muzzle. 

During the life of the gun the area near the origin 
of rifling was completely protected by being beneath 
the sleeve. Failure occurred beyond the end of the 
sleeve, although the rate of erosion was much slower 
than that observed in the gun steel bore. The use of 
the sleeve in a chromium-plated bore increased the 
velocity-life of the gun four times compared with gun 
steel without the sleeve, as shown in Table 1 and in 
Figure 6. 

32-3-4 Browning Machine Gun Tests 

The purpose of the tests with the Browning machine 
gun action was to observe the behavior of the sleeve 


Table 1. Increase of velocity-life of caliber .50 barrel 
obtained by use of Fisa protector with chromium-plated 
bore. 



Velocity life* 


(rounds) 

Gun steel bore—control 

90 

0.005 in. chromium-plated bore—no Fisa 

225 

0.005 in. chromium-plated bore—Fisa protector 

400 


* Number of rounds fired for a velocity loss of 200 fps. 



Figure 6. Velocity loss of caliber .50 barrels fired at 
an initial muzzle velocity of 3,500 fps: (1) unprotected 
gun steel barrel; (2) chromium-plated barrel; and (3) 
chromium-plated barrel with Fisa protector. (This fig¬ 
ure has appeared as Figure 23 in NDRC Report 
No. A-449.) 

used with banded artillery-type projectiles under con¬ 
ditions of rapid fire. A special barrel was used that 
had the same external contour as the standard cali¬ 
ber .50 aircraft machine gun barrel. The rifling was 
made 0.010-in. deep, the origin of rifling was modified 
to accommodate the sleeve, and the bore was plated 84 
with chromium 0.0038-in. thick at the breech end and 
0.0045-in. thick at the muzzle. 

In the final test of cadmium-plated sleeves the gun 
was fired with a standard charge of I MR powder to 
give a muzzle velocity of 2,800 fps. A control test was 
fired with a plain steel barrel and ball bullets, M2. 
Each barrel was fired 750 rounds in three groups of 
250 rounds, each group consisting of five 50-round 
bursts fired at 1-min intervals [the 5 X 50 (1) schedule 
as defined in Section 23.1.3] at a cyclic rate of 600 to 
640 rounds a minute. 

At the end of 750 rounds (three groups) the chro¬ 
mium-plated barrel used with the Fisa protectors 
showed no drop in velocity and no change in dimen¬ 
sions at the origin of rifling. The gun steel barrel 
after the same number of rounds showed a velocity 
drop of 139 fps and an increase in land diameter of 
0.0108 in. at the origin of rifling. Furthermore, some 
of the bullets from this barrel keyholed in each 250- 
round group, whereas none of the bullets fired from 
the barrel with Fisa protectors keyholed. 


CONFIDENTIAL 








614 


THE FISA PROTECTOR AND CHROMIUM-PLATED BORE 


Although slight tearing of the sleeves was observed 
during the test, this did not interfere with the func¬ 
tioning or performance of the gun. 

32.4 CALIBER .60 EXPERIMENTS 

Acaliber .60 barrel, Ordnance DrawingNo. 7160408, 
was modified for the sleeve designated as A-104shown 
in Figure 2. The barrel was electropolished and chro¬ 
mium-plated. 84 The sleeves were drawn from copper- 
coated, flat sheet steel and cadmium plated to reduce 
the friction between the band and the sleeve during 
the engraving operation. 

It was necessary to modify the caliber .60projectile, 
T32 (then undergoing test by the Ordnance Depart¬ 
ment), by reducing the diameter of the body back of 
the bourrelet by 0.008 in. This was done so that the 
diameter of the sleeve could be reduced in sections D 
and E (see Figure 3) and the lands could be deeper in 
the areas designated as D' and E' in Figure 4. This 
was necessary to obtain deeper engraving of the ro¬ 
tating band of the projectile. The components of the 
assembled round, the assembled round, and a fired 
round are shown in Figure 1. 

A total of 80 rounds was fired. All sleeves were ex¬ 
tracted from the chamber and there was no stretch¬ 
ing or tearing of the sleeves. The chromium plate 
failed along the bore on account of a poor bond, and 
firing was discontinued. 

32 5 37-MM FIRING TESTS 

A 37-mm gun tube, M3 was step rifled for the 
Fisa protector and chromium plated after electro¬ 
polishing. A thickness of 0.0055- to 0.0060-in. chro¬ 
mium plate was applied. The sleeves were drawn from 
copper-coated, flat sheet steel. In assembling the 
round the standard M51B1 and M51B2, projectiles 
were used with the 37-mm cartridge case, M16. Double¬ 
base powder (FNH-M2) was used in all the tests. 

Preliminary firing tests showed that a crimp was 
necessary to lock the sleeve to the cartridge case 
strongly enough to insure satisfactory extraction of 


the sleeve with the case after firing, in particular at 
low pressures (30,000 to 40,000 psi-copper). 

In order to reduce the coefficient of friction be¬ 
tween the sleeve and the band the sleeves were coated 
with various low-friction coatings. The following 
sleeves were tested: 

1. Cadmium-plated—as drawn. 

2. Cadmium-plated—annealed sleeves. 

3. Chromium-plated—as drawn. 

4. Parco-Lubrized—as drawn. 

5. Parco-Lubrized—annealed. 

6 . Steel—as drawn. 

7. Brass—machined. 

Results of tests fired at 50,000-psi pressure (copper) 
and a velocity of 2,900 fps showed (1) all the brass 
sleeves were torn; (2) the cadmium-plated and Parco- 
Lubrized sleeves performed better than the steel 
sleeves as drawn; (3) the annealed sleeves were no 
better than the sleeves as drawn; and (4) Parco- 
Lubrizing the sleeve surface was slightly better than 
cadmium plating it. 

It was hoped that the sleeve could be slipped over 
existing assembled rounds without any change to the 
ammunition. Because of the excessive interference 
between the sleeve and the band in the grooves, how¬ 
ever, a small percentage of the sleeves were torn dur¬ 
ing engraving. 

A total of 162 rounds was fired in the chromium- 
plated 37-mm barrel. Excessive erosion of the chro¬ 
mium plate occurred at the maximum height of the 
lands and in the grooves just beyond the end of the 
sleeve. This has been attributed to improper appli¬ 
cation of chromium plate. As mentioned in Chapter 
20 , there has not yet been sufficient experience in 
applying thick chromium plates to this size gun tube 
to give assurance of success in every plating job. 

It is recommended that further tests of 37-mm Fisa 
protectors be made, using a newly plated tube and 
projectiles having a reduced band diameter so that 
there is no interference in the grooves. The radial 
load might also be further reduced by increasing and 
deepening the cannelures, as described in Section 
27.4.2. 


CONFIDENTIAL 




Chapter 33 

PRACTICAL HYPERYELOCITY GUNS 

By L. H. Adams,* * J• S. Burlew, h and E. L. Rose c 


331 PRESENT STATUS OF THE 
DEVELOPMENT 

Understanding of the Erosion 

Process 

T he control of gun erosion has been the crux of the 
hypervelocity problem, as was brought out in 
Chapter 1. The studies of the mechanism of erosion 
conducted by Division 1 have led to an understanding 
of the part that different factors play in that process, 
as described in Chapter 13. Guided by this knowledge 
the search for erosion-resistant materials, described 
in Chapter 16, was focused on a few metals and their 
alloys. The need for fabricating these materials in 
forms suitable for application to gun bores led to a 
large number of metallurgical investigations, de¬ 
scribed in Chapters 17, 18, 19, and 20. 

Success in this development work has already been 
attained as far as stellite liners for machine gun bar¬ 
rels and chromium plating of such barrels are con¬ 
cerned, as is recounted in Chapters 22, 23, 24-, and 25. 
The application of these same materials to cannon, 
however, has not yet been realized. Because of the lar¬ 
ger size, the problem is considerably more difficult. 
The first attempt to apply chromium electroplate to 
a cannon tube in conjunction with pre-engraved pro¬ 
jectiles is described in Section 31.7. The preliminary 
efforts that have been made to apply stellite and 
molybdenum as liners for cannon are described in 
the next two sections of this chapter. 

331,2 Stellite Liners for Cannon 

Breech liners and linings of Stellite No. 21 have 
been applied experimentally to 37-mm gun tubes, but 
exhaustive tests have not been completed. Plans for 

a Chief, Division 1, NDRC. (Present address: Geophysical 

• Laboratory, Carnegie Institution of Washington.) Author of 
Section 33.3. 

b Technical Aide, Division 1, NDRC. (Present address: 
Geophysical Laboratory, Carnegie Institution of Washington.) 
Author of Section 33.1. 

c Member, Division 1, NDRC. (Present address: Consulting 
Engineer, Jones & Lamson Machine Company and Bryant 
Chucking Grinder Company, Springfield, Vermont.) Author 
of Section 33.2 


this work were started by Division 1 at the informal 
suggestion of the Ordnance Department, as soon as 
the first successful tests of a liner of Stellite No. 21 had 
been made in a caliber .50 aircraft machine gun bar¬ 
rel, as related in Section 22.2.2. At that time experi¬ 
ence in making large “investment castings” of Stel¬ 
lite No. 21 was so slight that preparation of liners was 
undertaken by other methods also. Because of delays 
in machining the liners made by the other methods, 
an investment-cast liner also was completed in time 
to be included in the group tested. 

Test Assemblies 

Two 37-mm gun tubes, M3, containing bore sur¬ 
faces of Stellite No. 21 for 20.8 in. forward from the 
origin of rifling were prepared. 80 One tube contained a 
flanged investment-cast liner of stellite shrunk di¬ 
rectly into the breech end of the tube and held in 
place by a steel breech nut containing a portion of the 
chamber. This design is essentially the same as that 
employed in the insertion of investment-cast liners 
for machine gun barrels, shown in Figure 1 of Chap¬ 
ter 22, except that in proportion to its caliber, the 
wall thickness of the 37-mm liner is less. 

The second tube was provided with two inter¬ 
changeable, replaceable steel liners, similar in design 
to the replaceable liners for the 90-mm gun described 
in Section 26.3. These two steel liners had been given 
a lining of stellite, applied to the bore surface by the 
process of progressive static infusion 89 described in 
Section 19.4.3. These two liners were prepared in the 
same way, except that one of them was heat-treated 
after the infusion process to increase the strength of 
the steel. 

Firing Tests 

The three liners were fired at the Ordnance Re¬ 
search Center, Aberdeen Proving Ground. 221 The in¬ 
vestment-cast liner and the static-infused, heat- 
treated liner were each fired 104 rounds with single¬ 
base powder. After 12 rounds for the establishment of 
charge and proof had been fired, the liners were fired 
at excess pressure (averaging 42,000 psi, copper) for 


CONFIDENTIAL 


615 



616 


PRACTICAL HYPERVELOCITY GUNS 


7 rounds, followed by a total of 85 rapid-fire rounds 
(6 rounds a minute), at an average muzzle velocity 
of 2,575 fps for the investment-cast liner, and 2,540 
fps for the other. After the first 19 rounds there was 
some expansion of the bore diameter, which was 
greatest at and near the origin of rifling. Since no sig¬ 
nificant changes in bore dimensions resulted from the 
firing of the remaining rounds, it was concluded that 
the early increase in diameter represented irreversible 
expansion of the liner, caused by both band pressure 
and powder pressure. As a result of a small decrease 
in liner length, an opening appeared at the breech end 
of the liner, permitting the cartridge case to expand 
into this opening, thus causing difficulties in extrac¬ 
tion. The opening was observed after 29 rounds had 
been fired and increased in width during the firing of 
the remaining 75 rounds. 

To determine whether melting of the stellite would 
take place, the investment-cast liner was then fired 
43 rounds with double-base powder (FNH, M2, which 
contained 20% nitroglycerin). In order to obtain a 
muzzle velocity of about 2,900 fps, the pressure was 
increased to 50,000 psi (copper). Eight proof rounds 
were followed by 10 rounds at an average rate of 4 
rounds a minute. A horoscope examination showed a 
marked change in the appearance of the bore surface 
at the origin of rifling, indicating that fusion of the 
surface had taken place. An additional 25 rounds (at 
a rate of about 6 a minute) were fired later. 

The third liner—a static-infused liner that had not 
been heat-treated after the infusion process—was to 
have been fired with single-base powder at a higher 
pressure than that used for the other two liners. There 
was difficulty in achieving the desired pressure with 
the powders available at the Proving Ground. After 
29 rounds had been fired, examination of the liner 
indicated that lack of heat treatment of the steel 
caused no difficulty. 

These firing tests demonstrated that stellite with¬ 
stands the action of single-base powder in a 37-mm 
gun fired at normal velocities, but that it is melted 
when double-base powder is fired at a slightly higher 
velocity. They also showed that the stellite is not 
cracked by the stresses of firing in this gun. Appli¬ 
cation of stellite to the bore surface by static infusion 
seems preferable to the use of an investment casting, 
because of a change in dimensions or movement of 
the liner, which opens a crack at the rear of the liner. 
These tests were not carried far enough with single¬ 
base powder to determine how much better Stellite 
No. 21 resists erosion in a 37-mm gun than does steel. 


Future Possibilities 

The surface melting of the 37-mm stellite liner 
when fired with double-base powder had been pre¬ 
dicted by calculations of the bore-surface temperature 
by the Hobstetter method described in Section 5.4.1. 
Calculations for other guns fired at hypervelocities, 
even with single-base powder, indicate that surface 
melting of stellite will occur in general in medium-and 
large-caliber guns. Hence stellite as such cannot be 
regarded as the complete solution to the hyperveloc¬ 
ity problem. 

One possible way of circumventing this difficulty 
would be to coat the stellite with a material of higher 
melting point, such as chromium or molybdenum. 
This possibility had been the motivation behind the 
efforts to plate molybdenum or chromium on stellite 
by the carbonyl process, as described in Sections 21.3 
and 21.6. Although success was not achieved, future 
discoveries in connection with the carbonyl process 
or some other plating process might make this possi¬ 
bility a reality. 

Even though the tests of 37-mm stellite liners did 
not result in a demonstration of wide applicability of 
this material to cannon, the investigation was well 
worth while. The experience gained in it is helpful in 
planning uses for other erosion-resistant materials. 
Thus the chromium-base alloys described in Chapter 
17 are potentially very important as bore-surface 
materials for cannon. As soon as a casting of suitable 
properties has been prepared, it will be a straight¬ 
forward job to make it into a liner for trial with as¬ 
surance that the design of the liner itself would be 
satisfactory. As mentioned in the next section, some 
of this experience has also been helpful in planning a 
molybdenum liner for a medium-caliber gun. 

331,3 Plans for 

Molybdenum Liner for 3-in. Gun 

The firing tests of caliber .50 molybdenum liners 
described in Section 18.6 showed that considerable 
progress had been made in improving the properties 
of molybdenum so that it would be suitable for use as 
a gun liner. The next step was to try a molybdenum 
liner in a medium-caliber gun, preferably of higher 
than conventional velocity. At the suggestion of the 
Bureau of Ordnance the Navy’s new 3-in./70-cal. gun 
was selected for this purpose in the spring of 1945. 
This gun, just then in process of development, was 
designed as a high-velocity weapon firing automati- 


CONFIDENTIAL 



PRESENT STATUS OF THE DEVELOPMENT 


617 


cally for antiaircraft use. The muzzle velocity is 3,400 
fps at a maximum true pressure of 60,000 psi. The 
gun is intended to be fired at a rate of 90 rounds a 
minute for as long as one minute. The heating of the 
barrel under these conditions is discussed in Section 
5.5.3. 

Liner Designs 

Arrangements were made to expand the facilities of 
the Westinghouse Lamp Division to make possible 
the fabrication of larger pieces of molybdenum, as 
described in Section 18.3.4. Then a planning con¬ 
ference was held in the office of Division 1 on August 
15, 1945 (the day after V-J Day), at which a number 
of Division 1 contractors d considered various means 
of inserting liners of both molybdenum and chrom¬ 
ium-base alloy in the 3-in./70-cal. gun. 

One design of molybdenum liner proposed was simi¬ 
lar to the investment-cast stellite liners. The molyb¬ 
denum, which might be fabricated in several staves 
(Section 18.5.2) or as a seamless tube (Section 18.5.3), 
would be shrunk into the recessed gun tube and then 
held in place by a short steel chamber liner. 

A simpler way of handling a multistave liner would 
be to shrink the staves into a steel carrier and then 
insert this composite liner into the recessed gun tube, 
in the . same way the composite stellite liner was in¬ 
serted in a 37-mm tube (Section 33.1.2). 

Both the foregoing designs would involve recessing 
the gun tube from the breech end, which would make 
it necessary also to line the chamber. A way of avoid¬ 
ing this is to make the gun tube in two longitudinal 
segments held together by a locking ring. The rear 
segment, which contains the chamber, is recessed 
rearward from its front end to receive the liner. The 
front segment is turned down to fit into the forward 
end of this recess and make a seal with the front end 
of the liner. This design is a modification of those 
used by the Geophysical Laboratory (Section 11.2.2) 
and by The Franklin Institute (Section 11.2.1) in 
caliber .50 guns for testing materials for erosion 
resistance. Another advantage of this design is that 
by using a relatively short muzzle section it is simpler 
to chromium-plate the steel bore surface ahead of the 
molybdenum liner. 

d No formal report of this conference has been issued. It was 
attended by representatives of the following contractors in 
addition to Division 1 and Navy Bureau of Ordnance person¬ 
nel : Crane Company, The Franklin Institute, Un ion Carbide and 
Carbon Research Laboratories, Westinghouse Lamp Division, 
and Westinghouse Research Laboratories. 


Continuation of Project 

After termination of Division l’s contracts the 
Navy Bureau of Ordnance contracted with the West¬ 
inghouse Electric Corporation to continue the devel¬ 
opment of molybdenum liners. Following further 
tests of caliber .50 models, using molybdenum fabri¬ 
cated with the enlarged facilities mentioned in Section 
18.3.4, it is planned to make preliminary tests with a 
40-mm liner of the two-segment design described in 
the previous paragraph. The first liners will be two- 
stave ones, but it is hoped that eventually tube-mak¬ 
ing experiments will result in a successful seamless 
liner. Eventually it may be possible to utilize molyb¬ 
denum in the form of a seamless liner the whole length 
of the gun, as shown in Figure 1. This design might 
be the most economical. 

331,4 Plans for A-Z Gun 

As a culmination of Division l’s investigation of 
erosion-resistant materials, the A-Z project was under¬ 
taken, whose ultimate purpose was the development 
of a gun tube for an entirely new weapon which would 
embody the best features of erosion control developed 
by the Division. It was hoped that the project would 
result in a gun whose velocity was the highest practi¬ 
cable. By the introduction of liners or other means of 
erosion control, surface cracking in the structural por¬ 
tion of the gun tube would be eliminated. The way 
would then be open to work the gun materials at a 
higher stress and to use materials of higher strength 
and lower elongation than is permissible in existing 
weapons. The gun was to be designed to fire both 
high-explosive and armor-piercing projectiles, so that 
a decision as to its eventual application could be made 
later. 

As a first step, plans were formulated for an experi¬ 
mental, medium-caliber hypervelocity gun which 
could be used as a testing gun for three types of ex¬ 
periments : insertion of an erosion-resistant bore-sur¬ 
face material; trial of a high-strength gun steel; and 
ignition of powder at high densities of loading, pos¬ 
sibly leading to the development of a higher density 
powder grain. 

The design of a 90-mm gun, having a muzzle ve¬ 
locity of 4,200 fps and a maximum powder pressure of 
60,000 psi, was decided upon. The basic specifications 
for the gun are given in Table 12 of Chapter 31. The 
first trial was to have been with a chromium-plated 
tube, fired with pre-engraved projectiles (Section 


CONFIDENTIAL 




618 


PRACTICAL HYPERVELOCITY GUNS 



The hypervelocity gun of the future may well be one containing a full-length liner of hardened molybdenum 


Figure 1. 
alloy. 

31.8). Additional rifled tubes, counterboredto receive 
liners, were to be prepared so that as erosion-resistant 
materials became available their performance as 
breech liners could be evaluated. When Division l’s 
investigations were terminated, tube forgings for the 
A-Z gun were being prepared. Two of them were com¬ 
pleted to the stage of smooth-boring and chambering, 
and were turned over to the Army Ordnance Depart¬ 
ment for experimental use. 


33 2 EVALUATION OF MILITARY 
IMPORTANCE OF HYPERVELOCITY 
GUNS FIRING SUBCALIBER 
PROJECTILES 

33 2 1 General Observations 

Military Economics 

Many arguments favor the use of hypervelocity 
guns, but verbal argumentation leads to no conclu¬ 
sions regarding the relative merits of military weapons. 
This is partly because of the many points of view from 
which the problem of evaluation can be approached. 
There appear to be two extreme approaches between 
which the many alternatives lie. 

In the military emergency where failure means in¬ 
calculable loss, the cost of attack or defense, however 
figured, is no consideration. The sole questions of 
importance are what to do and what is within reach 
to do with. It is too late for planning and economic 
study to contribute to the solution and the only argu¬ 
ment which avails anything comes as the post-mor¬ 
tem where successes and errors are unearthed, dis¬ 


cussed, and recorded for the benefit of future gener¬ 
ations. 

In broad-gauge military planning, during times of 
peace, the evaluation of weapons is an economic 
problem. It is concerned both with the economics 
of production in time of peace, and with the economics 
of production and use in time of war. In time of 
peace the economy is a dollar economy; in time of 
war it is controlled by the use of critical labor and 
critical material. The labor includes all the labor from 
the harvesting of raw materials, through manufacture 
of materiel, manufacture of transport equipment, 
labor of transport, labor of use, and labor of mainte¬ 
nance and repair. Little can be accomplished in the 
attempt to design equipment for the unpredictable 
emergency. Much can be accomplished toward re¬ 
ducing the probability of the emergency by advance 
planning of the economical conduct of war so as to 
make available manpower and materials most effec¬ 
tive for the purpose. 

The most effective weapon is the one which in¬ 
volves the minimum cost for a given damage to the 
enemy. The relative evaluation of cost and damage 
is involved in questions of military economics and 
strategy. In general, the problem is beyond the scope 
of this report. However, when the destructiveness of 
a given weapon can be increased many fold without 
additional cost, there is no question regarding its 
economics, military or otherwise. It is from this stand¬ 
point that we approach the evaluation of guns of 
established design, firing subcaliber projectiles at 
hypervelocity. Such projectiles include both sabot- 
projectiles (Chapter 29) and skirted projectiles 
(Chapter 30). 


CONFIDENTIAL 





MILITARY IMPORTANCE OF SUBCALIBER PROJECTILES 


619 


The evaluation of the actual overall military econ¬ 
omy of this accomplishment is a matter deserving 
further careful and extended study. It is to be hoped 
that the simple approach to the problem outlined 
here may point the way for the conduct of more 
elaborate studies of weapons in general. 

Pros and Cons of Subcaliber Projectiles 

Many valuable weapons have been neglected be¬ 
cause, in the absence of simple methods for analytical 
evaluation, resort has been to opinion or to limited or 
inconclusive experience in judging their usefulness. 
The use of subcaliber projectiles in standard-bore 
guns has suffered this fate. It has long been known 
that muzzle velocities could be increased by use of 
subcaliber projectiles. 190 But it has also been known 
that reducing the caliber of projectiles reduces the 
effective radius of burst of the high-explosive projec¬ 
tile and the relative armor penetration of the armor¬ 
piercing projectile. It has also been known that a 
reduction in caliber increases the retardation of the 
projectile and therefore its loss of velocity over a 
given range. 

These arguments have been aimed against the use 
of the subcaliber projectile without any attempt at a 
quantitative determination of the point of diminish¬ 
ing returns, the point at which the sacrifices offset the 
gains. The fact that there is a considerable region 
within which the net gains are substantial in spite of 
apparent disadvantages has been completely passed 
over. The fundamental reason behind this has been 
the fact that the method of handling the ballistic 
computation necessary in evaluating guns has been 
cumbersome and the computations too laborious for 
the extensive studies necessary. 

The methods outlined in this report are not exact. 
However, their application to problems with known 
solutions shows results within an accuracy of a few 
per cent. Furthermore, the accuracy is completely 
obscured by the manyfold gains indicated so that 
there is no question as to the validity of conclusions 
except, possibly, under extreme conditions which lie 
outside of the present region of practical importance. 
The same methods could undoubtedly be applied to 
many other problems concerned with the evaluation 
of guns. 

Increased Projectile Velocity 

Raising the velocity of the projectile fired from a 


gun increases the effectiveness of the gun in one or 
more of several ways, depending on the use. These 
may be summarized as follows. 

1 . In unaided (no director) fire aimed at a stationary 
or moving target, it increases the allowable error in 
estimating range without reducing the probability of 
hitting. 

2. In unaided fire aimed at a moving target, it 
reduces the necessary lead angle. Since the allowable 
absolute error in lead angle is independent of the 
lead and since the absolute accuracy of estimating 
lead is better the smaller the lead, the probability 
of hitting is improved. 

3. In director-controlled fire against a maneuver¬ 
ing target, the probability of hitting increases more 
rapidly than the velocity 164 since the region into 
which the target can maneuver in the time of flight 
is diminished in volume. The probability of hitting 
is greatly increased by attaining an increase in 
velocity. 

4. The thickness of armor that can be penetrated 
at a given range increases with velocity up to a criti¬ 
cal velocity 205 at which the projectile breaks up; and 
the range at which a given thickness can be pene¬ 
trated increases with velocity. 

When subcaliber ammunition is used to increase 
velocity, there is a net gain in effectiveness in spite of 
the reduced relative effectiveness of the smaller pro¬ 
jectile. The overall effectiveness may be considered 
as the product of the probability of a damaging hit 
and the measure of effectiveness of the hit. As the size 
of the subcaliber projectile is varied downward this 
product, for each situation, increases for some range 
of caliber reduction, reaches a maximum, and then 
diminishes. The only situations in which this is not 
true are those in which the probability of hitting is 
not affected by velocity (for instance in director-con¬ 
trolled fire against a fixed target or against a target 
moving on a definitely fixed and predictable course); 
or those in which the power of the standard projectile 
is just sufficient to destroy the target and any reduc¬ 
tion of caliber decreases the destructive effect. 

In attacks against armor, the effectiveness of a 
projectile which has insufficient energy to penetrate, 
is always increased by the use of a subcaliber projec¬ 
tile, provided that (1) the velocity required for pene¬ 
tration by the subcaliber projectile is not above its 
shatter value (see Section 33.2.4), and (2) the range 
is not so great that air-resistance velocity attenuation 
wipes out the potential gain. In antiaircraft use, 
where the high-explosive projectile has a finite range 


CONFIDENTIAL 




620 


PRACTICAL HYPERVELOCITY GUNS 


of effectiveness of fragmentation, the gains by the use 
of subcaliber projectiles are always very striking, over 
a large range of target distances and with standard- 
caliber velocities as high as 6,000 fps. It does not ap¬ 
pear that any velocity for the standard-caliber pro¬ 
jectile attainable with a practicable gun design is so 
high that the effectiveness of the gun in antiaircraft 
service may not be materially improved by the use of 
a subcaliber projectile. 

There are three methods for increasing the muzzle 
velocity for a given gun and powder charge. 

1 . Use of lightweight, standard-caliber projectile. 

2. Use of a subcaliber projectile with discarding 
sabot in a standard-bore gun. (See Chapter 29.) 

3. Use of a subcaliber, skirted projectile in a 
standard-bore gun fitted with a tapered muzzle 
adaptor, or in a one-piece tapered-bore gun. (See 
Chapter 30.) 

Both types of subcaliber projectiles have a con¬ 
siderable advantage over the standard projectile of 
equal weight in that they suffer less velocity loss for 
a given muzzle velocity, over a given range. The 
skirted projectile has a definite advantage over the 
sabot-type projectile of the same weight, in that the 
entire initial momentum of the skirted projectile is 
available for sustaining velocity against air resist¬ 
ance, whereas the momentum of the sabot is lost at 
the muzzle of the gun. This fact shows up quite 
strikingly in some of the comparisons which follow. 
Furthermore, since the skirted projectile fits the bore 
tightly throughout its travel, it leaves the bore with 
less yaw and its accuracy is measurably superior. 

Tungsten Carbide Cores 

The development of the use of tungsten carbide 
cores in armor-piercing projectiles has gone hand-in- 
hand with the development of subcaliber projectiles 
and guns for firing them. This is not because of any 
unique advantage deriving from the combination, for 
the use of carbide cores in standard-caliber projec¬ 
tiles possesses similar advantages. However, since the 
purpose of the subcaliber development has been the 
attainment of the utmost in penetration, the adoption 
of the tungsten carbide core has been a natural ac¬ 
companiment. 

The gain from the use of tungsten carbide cores 
derives particularly from the high density of the 
material, which increases the energy density over the 
impact area, with a given velocity and from the high 
hardness which keeps down the deformation of the 


projectile on impact. The gain in penetration is also 
influenced, probably in a minor way, by the low coef¬ 
ficient of friction of tungsten carbide on steel. Fur¬ 
thermore, tungsten carbide of suitable cobalt content 
shows some increase in shatter velocity over steel. 

All these factors combine to produce a very sub¬ 
stantial gain in armor penetration through the use of 
tungsten carbide, and future attempts to develop 
projectiles of increased armor penetration should in¬ 
clude a concentrated effort to improve the methods of 
using tungsten carbide. Such efforts should, in part, 
be directed at the development of a tungsten carbide 
head for a high-explosive, armor-piercing projectile. 
When used as a subcaliber projectile in a hyperve- 
locity gun, such a projectile should have great effec¬ 
tiveness. 

33 2 2 Velocity Gains From the Use of 
Subcaliber Projectiles 

Limiting Effect of Powder Gas Kinetic Energy 

Several years ago Langweiler fired an 8-mm/125- 
caliber gun with a fixed charge of pulverized nitro¬ 
cellulose powder and with projectiles varying in 
weight from 1.18 times the charge weight to l /u the 
charge weight. Results of these firings 20498 show quite 
strikingly the effect of kinetic energy of the powder 
gas in limiting the velocity of the projectile. In partic¬ 
ular they show that the velocity-charge function 
can be estimated with good accuracy by assuming 
approximately one-fourth of the charge weight added 
to the projectile weight and treating the kinetic ener¬ 
gy of this combination as proportional to the powder 
energy. [See equation (4) in Section 3.2.2.] 

Figure 2 is a graph of the Langweiler values of the 
velocity plotted against ratios of charge to projectile 
mass. The curve was drawn through points calculated 
from equation (1), 



in which Uis the calculated velocity, M is the projec¬ 
tile mass, and C is mass of charge. The assumption is 
made that a fraction 1/a of the powder gas mass 
moves with the projectile. Vk is a constant with the 
dimensions of velocity, equal to V when M/C = 
(1 - 1/a). 

The value of a was determined from the experi¬ 
mental data in Table 1. In this table, also, calculated 


CONFIDENTIAL 




MILITARY IMPORTANCE OF SUBCALIBER PROJECTILES 


621 


Table 1. Experimental data from firing of 8-mm/125- 
cal. gun. 498 


C/M 

F exp 

fps 

F calc 

fps 

AF 

per cent 

0.85 

3,970 

4,000 

0.8 

3.20 

6,365 

6,365 

0 

5.80 

7,380 

7,480 

1.4 

11.00 

7,970 

8,280 

4.2 

22.00 

8,950 

8,900 

0.6 

44.00 

9,150 

9,150 

0 


values of velocity are compared with experimental 
values. Fitting the curve to the points where C/M 
= 44 and C/M = 3.2, a is found to be 0.247 and V 0 is 
4,780 fps. The approximation is quite good enough 
for our purpose. 

Also plotted on Figure 2 are points of velocity 
versus C/M from data obtained at Carderock with 


the 3-in./50-cal. Navy gun (Section 4.2.2). These 
points are seen to fit the curve almost perfectly. Also 
plotted are a number of points obtained with a large 
variety of projectile designs and with various powder 
grain sizes, from firings with the 57/40-mm tapered- 
bore gun with skirted projectiles (Chapter 30). It is 
to be noted that for low values of C/M these latter 
points tend to lie below the Langweiler curve while 
for higher values they tend to lie above it. However, 
the highest value attained lies on the curve. There 
are several reasons for this behavior. 

1 . The friction in the tapered-bore gun is high 
and tends to displace the point of zero velocity to the 
right. 

2 . The friction at low velocities is substantially 
constant and it becomes less important as the powder 
charge increases so the two curves approach each 
other. 



Figure 2. Velocity vs ratio of charge to projectile. 


CONFIDENTIAL 



































622 


PRACTICAL HYPERVELOCITY GUNS 


3. The Langweiler curve was taken with constant 
charge while the 57/40-mm data were taken with 
varying charge and nearly fixed projectile mass. 

4. As the velocity of the projectile increases, sur¬ 
face friction apparently begins to melt the projectile 
surface material and the lubricating effect of this 
molten material reduces the friction; all indications 
are that the coefficient of friction at 4,000 fps is less 
than 1 per cent. 

5. The relative heat loss, from powder gas to walls, 
decreases as the size of the gun increases. The thermal 
efficiency of the larger gun is therefore higher. 

6. The horizontal spread of points is due to varia¬ 
tion of powder temperature and to variation in pro¬ 
jectile friction from one design to another. Gun wear 
also contributes to this spread. 

The improvement in efficiency for higher charge 
and the greater efficiency of the larger gun cause the 
57/40-mm curve to cross the Langweiler curve. The 
57-mm standard load is a 6.28-lb projectile propelled 
by 36.5 oz of powder. The muzzle velocity is 2,800 
fps and the charge-projectile mass ratio is 0.37. This 
point lies on the Langweiler curve as do skirted pro¬ 
jectiles fired with 36.5 oz of powder. It is therefore 
clear that the skirted projectiles fired from a tapered- 
bore gun closely follow the performance indicated by 
Langweiler’s curve. The velocity performance of the 
tapered-bore gun or of the gun firing a sabot-projec¬ 
tile can be computed from equation (2) 



1+ 0.25(C/ITo) 1 1/2 

W a /Wo + 0.25(C/TF 0 )J 


(2) 


in which Vo is the velocity for the standard gun and 
C is the standard charge, for the standard projectile 
of weight Wo. V 8 is the velocity that would be ob¬ 
tained from the subcaliber projectile of weight W s . If 
the charge is to be varied, additional data are neces¬ 
sary to introduce the effect of friction and cooling of 
the powder gas. Equation (2) provides a simple ex¬ 
pression for use in determining subcaliber projectile 
velocities for the purpose of approximate analyses. 


Effect of Gun Caliber on Muzzle Velocity 

• 

When an homologous series of projectile designs has 
been prepared the weights of the projectiles will vary 
with the cube of the caliber. However, when subcali¬ 
ber projectiles are fired, additional parts are necessary 
to fill the bore, serving to support the powder pressure 
and to guide the projectile down the bore. These parts 
add weight. A rough analysis indicates that the gross 


weight of a subcaliber projectile can be related to the 
weight of the standard projectile by equation (3) 


+ o(l - p) 2 ( 1 + P + (3) 

in which W a is the weight of the subcaliber projectile, 
Wo is the weight of the standard projectile, p is the 
ratio of subcaliber to standard caliber, and <r and 0 
are design constants. For a skirted projectile, <r is 
approximately 0.16 and 0 = 0. For sabot-type pro¬ 
jectiles, a is substantially the same and 0 = 3. 

By combining equations (2) and (3), the muzzle 
velocity V ms of a subcaliber projectile is given by 
equation (4). 

V m8 = Vo • 

1 + 0.25(C/IFo) 

p 3 + o-(l - p) 2 (l + p + 0) + 0.25(C/TFo) 

Effect of Range on Velocity and Time 
of Flight of Subcaliber Projectiles 

As indicated above, the analytical treatment of the 
relative effectiveness of different types and sizes of 
guns has always been complicated by the traditional 
empirical handling of the ballistic aspects of the prob¬ 
lem. In general it has not been practicable to effect a 
valid comparison without constructing and firing 
actual full-sized models. It is quite obvious that this 
fact has had an adverse effect on the cost of develop¬ 
ing ordnance and on the interest taken in the con¬ 
sideration of the potentialities of new weapons. 

Fortunately, when the field of hypervelocity is 
entered, the situation becomes greatly simplified. In 
unpublished studies of ballistic tables it has been 
found that the ballistic performance of projectiles of 
good aerodyamic design can be treated with sufficient 
accuracy for all purposes except actual fire control in 
service, by very simple functions, for a considerable 
range of velocities above that of sound. For guns with 
high velocities the accuracy is good throughout the 
range of target distances for which tables are generally 
computed. 

Over a considerable range of velocities above that 
of sound, with good aerodynamic projectile design, 
the resistance force varies closely with the three- 
halves power of the velocity. In Chapter 8 closer ap¬ 
proximations effective over a limited range of velocity 
are indicated. The three-halves power law is quite 
sufficiently accurate for many purposes up to Mach 
numbers approaching 10. 



CONFIDENTIAL 






MILITARY IMPORTANCE OF SUBCALIBER PROJECTILES 


623 


When the muzzle velocity is above, say, 2,600 fps, 
the range tables show a maximum variation of range 
for constant time of flight, with varying elevation, of 
only a few per cent. In the case of the 120-mm gun, 
Ml, this variation is of the order of only 1 per cent. 
This fact is due to the compensating effects of re¬ 
duced air resistance at high altitudes in balancing 
loss of velocity due to gravity. In view of these facts 
very good analytical approximations can be made by 
neglecting altitude-density effects and the effect of 
gravity on the flight of the projectile. Under these 
circumstances the differential equation of projectile 
motion along the line of sight may be written in the 
form of equation (5), 



in which r is the distance along the line of sight, t is 
the time of flight, and K is the aerodynamic retarda¬ 
tion constant. The solution of equation (5) may be 
written in the forms (6), (7), and (8), 


ut 

(6) 

r - Ky/Yd ’ 

+ 2 

, r/u 

i 

2 y/u’ 

(7) 

( Kr \ 2 

(8) 

F ' = ’V - 2 Vu) ’ 


fit its ballistic table quite well. For instance, for 25 
sec time of flight the range on a horizontal line of 
sight is 14,200 yd. For these conditions calculation 
shows^that u = 3,000 fps; Kr/2^u = 0.432; 
K/2\/u = 0.0000304; K = 0.00191. Calculations of 
i by means of equation (7) deviate only slightly 
(between 1 and 2 per cent) from corresponding values 
of t given in the range tables. 

When the projectile size is varied while maintain¬ 
ing constant aerodynamical design, K varies directly 
with the cross-sectional area of the projectile and in¬ 
versely with its mass. Furthermore K depends pri¬ 
marily on the nose design and changes in length 
affect it little. For subcaliber projectiles we may there¬ 
fore write equation (9), 


K = K 0 ^ = A'o 


V + ff(l - p) 2 (l + p + 0) 


(9) 


in which Ko and K are the constants for the standard 
and subcaliber projectiles respectively. 

In using this relation we must bear in mind that, 
in the case of the sabot-projectile, the sabot parts are 
discarded at the muzzle. Hence a and 0 become zero 
for the sabot-projectile in flight. In the case of the 
skirted projectile, /3 isO. We therefore obtain equations 
(10) and (11). 


K (skirted) 
K (sabot) 


K v + <r(l -p) 2 ( 1 + pY 



P 


( 10 ) 

(ID 


where u is the initial velocity of the projectile, and 
Vr is velocity at range r. 

The 120-mm gun, Ml, is a high-velocity (3,150 fps) 
gun with a good projectile design. Ballistic computa¬ 
tions based on the three-halves power resistance law 


By making use of equations (4) to (11) we may now 
write, for the velocity at any range, equations (12) 
and (13). 

For values of p close to unity, equations (12) and 
(13) can be approximated by equation (14). 


T7 _ T / f 1+0.25(C/JFo) V» 

Vr (skirted) Vo + _ p)!(1 + p ) + 0.25 (C/Wo) J ' 


1 - 


KopV + ff(l - p) 2 (1 + p) + 0.25(C/Tro)]w ) 2 

2V%[1 + 0.25(C/JF 0 )] 1/4 [p 3 + <r(l - p) 2 (1 + p)]J ■ 


v y ( _ 1 + 0.25(C/Wo) _ V' 2 

r <sabot) ° V + <K 1 - P) 2 (l + P + /3) + 0.25(C/IFo)J ' 


i K n[p 3 + g (l - p) 2 (l + p + 0 ) + 0.25(C/FTo)] 1 ' 4 ) 2 


V r 


(skirted) 


_ v =v v [ 1+0.25(C/W 0 ) > 
’ T (sabot) V r L p 3 + 0.25( C/W 7 ) J 


2p V / t 7 o[l + 0.25(C/Wo)] 1/4 

Ko 


1 - 


2p\/Vo 


-[p 3 + 0.25(C/T^o)] I/4 


( 12 ) 


(13) 

(14) 


CONFIDENTIAL 














624 


PRACTICAL HYPERVELOCITY GUNS 


It is apparent from inspection of this relation that 
while decreasing the projectile caliber always reduces 
the range at which the velocity reaches zero, there is still 
always a finite range of target distance within which 
striking velocity is increased by reducing caliber. 

Gains in Effectiveness of Fire 
against Moving Targets Resulting 
from Use of Subcaliber Projectiles 

Reduction of Time of Flight 

From equation (7) and the expressions for K and 
W/W 0 , the time of flight t of a subcaliber projectile 
is found to be given by equations (15) and (16). 

For a standard diameter projectile of reduced 
weight equal to that of the skirted projectile, the time 
of flight is given by equation (17). 


Table 2. Ratios of times of flight for various types of 
projectiles. 


p 

10,000-yd range 
td/to ts/to tl/to 

15,000-yd range 

td/to ts/to U/to 

1.0 

1.00 

1.00 

1.00 

1.00 

1.00 

1.00 

0.9 

0.90 

0.90 

0.92 

0.90 

0.91 

1.12 

0.8 

0.76 

0.81 

1.06 

0.81 

0.82 

1.60 

0.7 

0.70 

0.73 

1.37 

0.69 

0.82 

00 

0.6 

0.59 

0.66 

2.30 

0.60 

0.89 


0.5 

0.52 

0.65 

00 

0.61 

1.20 


0.4 

0.44 

0.78 


0.72 

3.50 



power of the projectile itself. It is important to con¬ 
sider the situation from the standpoint of overall 
gain. The probability of a direct hit is greatly in¬ 
creased with the subcaliber projectile, but the damage 
depends on the power of the hit. When the size of a 
high-explosive shell is reduced, its radius of effective¬ 
ness may be considered to be proportional to its diam- 


r [p 3 + <K 1 ~ p) 2 ( 1 + p) + 0.25(C/!Fo)] 1/2 _ 

ta (skirted) - y , . ,,/j j. Apr p' 2 [p 3 + <r(l ~ p) 2 (l + p) + 0.25 (C/IFq)] 1/4 1 

[1 + 0.25 (C/Wo)]' 1 {1 - 2V y^ (p z + „[1 - p]«[i + P p/-[ 1 + (0.25C/TFW 

r [p 3 + q~(l — p) 2 (1 + p + ft) + 0.25(C/fFo)] 1/2 __ # 

f Kqt f p 3 —|— cr( 1 — p) 2 (l + p + 0) + 0.25(C/TFo)] l/4 | 

Fp [1 + 0.25(C/TFp)] |l - —= P [1 + 0.2^C/Wo)] lli 1 


ts (sabot) 


tl (lightweight) 


|> 3 + g(l - p) 2 (1 + p) + 0.25( C/Wo)] 11 * 


tixAwr/mwn Apr [P 3 + 0.25(1 - p) 3 (i + p) + 0.25(C/W OP'*' 
[1 + 0.25(C/TF.)] [1 - 2 ^- T - [J + 0.25(C/fFo)] 1/4 [p 3 + <r(l - p) 2 U + p)] 


(15) 


(16) 


(17) 


For purposes of numerical comparison the 120-mm 
gun, Ml, has been chosen as it has a good projectile 
design and the normal muzzle velocity is high enough 
so that the approximate analysis applies throughout 
the range of uses. For this gun C/4IF 0 has the value 
0 .11. T a ble 2 shows a comparison of times of flight for 
10 ,000-vd range and 15,000-yd range of skirted and 
sabot-projectiles fired from this gun for varying val¬ 
ues of p. It is not practicable to reduce p below 0.5 and 
values of 0.6 and 0.7 are more practical. The value of 
0.4 has been included only to show the trend. There 
have also been included columns of time of flight t for 
lightweight, full-caliber projectiles having the same 
weight as the skirted projectile. The values shown in 
the table are the ratios of the time of flight of the 
reduced projectile to that (to) of the standard pro¬ 
jectile. 

Table 2 shows a substantial reduction in time of 
flight through the use of a subcaliber projectile. How¬ 
ever, this is at the cost of a reduction in the damage 


eter down to the point where the fragments become 
too small to effect critical damage. 

The superior performance of the skirted projectile 
results from the fact that no mass is discarded so that 
all the initial momentum is available for overcoming 
resistance. It is to be noted that the relative perform¬ 
ance of either subcaliber projectile is not greatly 
affected by range for values of p of 0.6 and above. The 
effect will, however, become more pronounced at 
greater ranges. The standard-caliber lightweight pro¬ 
jectile is obviously of little value except at relatively 
short range. 

Effect of Subcaliber on Probability of 
Damaging Hit 

The probability of an effective hit is a function of 
the number of shells which must be thrown into the 
space of future positions of the target to insure 
such a hit. This in turn depends on the dimensions of 


CONFIDENTIAL 






















MILITARY IMPORTANCE OF SUBCALIBER PROJECTILES 


625 


the vulnerable portion of the target and the radius of 
effectiveness of the burst. Numerically it is the re¬ 
ciprocal of the number of bursts spaced on some regular 
pattern of minimum density, sufficient for the pur¬ 
pose, modified by a coefficient which introduces the 
statistical aspects of the problem. For simplicity, a 
fair approximation can be made by treating both the 
target and the burst as spheres of radii R t and R b 
respectively. There are good reasons why this gives 
substantially as good an approximation as a more 
complicated analysis. If the bursts are placed at the 
corners of a cubical lattice having diagonal length of 
2{R t + Rb) and if V t is the volume of future positions 
of the target, the bursts per hit is expressed by 
qeuation (18), 


N h = 


3F* 


8 s(Rt T RbY 
The probability of a hit is given by equation (19). 


(18) 


^8 s(R t + RbY 

h 3 y 


(19) 


With a target subject to three-dimensional accelera¬ 
tions and moving at speed and with radial accelera¬ 
tion such that the change of course is 90 degrees or 
less in the time of flight of the projectile, the volume 
V t is approximately proportional to the sixth power 
of the time of flight t. The higher the velocity of the 
target the more closely this condition is met. We may 
therefore write equation (20). 

Pk, = 3 (20) 


Here P 0 is a statistical coefficient, 1/a is the ratio of 
the radius of a normal burst to that of the target, and 
R b is the ratio of the diameter of a subcaliber projec¬ 
tile to that of the normal projectile. 

The relative probability of a damaging hit with a 
subcaliber projectile as compared with that of a 
normal projectile is given by equation (21). 


Pk* _ (a + P ) 3 / t n V 
Vr Pkn ( a+iy\tj 


( 21 ) 


It is clear from this equation that the probability of 
a damaging hit is highest for the projectile with the 
shortest time of flight when p and a are constant. An 
analysis of the data shows that of the various forms 
of lightweight projectiles of equal power, fired from 
the same gun, the skirted projectile has the shortest 
time of flight. In an attack against a maneuvering 
target, there are very great advantages in the use of 


subcaliber projectiles, particularly those of the skirted 
type, in spite of the effect of reduced caliber in di¬ 
minishing the radius of effectiveness of the high- 
explosive burst. For the sabot-projectile the relative 
effectiveness at p = 0.6, near the point of greatest 
probability, is only about one-half that of the skirted 
projectile. This advantage of subcaliber projectiles 
persists into the region where practical considera¬ 
tions prevents further reduction of projectile cali¬ 
ber. 

The design of a deformable, high-explosive projec¬ 
tile presents certain difficulties which become more 
serious with decreasing values of p. Although the 
theoretical probability of damage is maximum in the 
neighborhood of p = 0.6 for a range of 10,000 yd, and 
for lower values of p at shorter ranges, the region 
p = 0.7 top = 0.8 is that to which practical consider¬ 
ations limit design. There is, however, an obviously 
substantial advantage to be gained even though we 
are limited to working in this region. Figure 3 il¬ 
lustrates graphically the advantages of subcaliber 
projectiles. 

While present design considerations tend to re¬ 
strict the caliber ratio to values of 0.7 and 0.8, Figure 
3 shows the great theoretical advantages to be gained 
by further decreasing p, down to values in the 
neighborhood of 0.55. In fact, in the least favorable 



0.2 0.4 0.6 0.8 i.o 


CALIBER RATIO 

Figure 3. Theoretical advantages of subcaliber pro¬ 
jectiles in increasing probability of damaging hit. 


CONFIDENTIAL 



















626 


PRACTICAL HYPERVELOCITY GUNS 


ease (a = 0), there is a gain of the order of 2 to 1 in 
going from p = 0.7 to p = 0.55. This fact creates a 
great opportunity for further gain by efforts to bridge 
this gap in design. The ultimate lower limit is that 
prescribed by the space available in the emergent 
bore for accommodation of the deformed rear skirt. 
For powder pressures in the neighborhood of 45,000 
psi, this limit is approximately at p = 0.5. For higher 
pressures the value is larger. The use of higher 
strength material is indicated. This, however, intro¬ 
duces a wear problem and demands development of 
wear-resistant liners. In the case of high-explosive 
projectiles, the weight of the rear skirt becomes im¬ 
portant in influencing the stability of flight. For 
designs developed to date it would probably be diffi¬ 
cult to maintain stability for p less than 0.75, except 
by greatly increased rate of spin. However, this mat¬ 
ter has not yet been thoroughly analyzed. 

For small values of p an excessive portion of the 
available space is taken up by the front skirt. When 
this occurs the space available for the high explosive 
is insufficient and the fragmentation of the projectile 
is apt to be unsatisfactory. New studies may lead to 
a better method for supporting the front end of the 
projectile. Even in the face of these difficulties the 
gains to be made by extending the range of p toward 
0.5 are such that even though the relative effective¬ 
ness be substantially reduced by the sacrifice of good 
design there may still be a net gain in increased 
probability of damage. This aspect of the design 
study is worthy of exhaustive treatment. Further 
consideration needs also to be given to effect of size 
on damage and the point at which the fragmentation 
is such as to produce ineffective bursts against the 
target. 

In the above analysis the standard powder case 
and charge are assumed, with only the projectile 
weight varying. A gain in the probability of a dam¬ 
aging hit is indicated throughout the range of practi¬ 
cable caliber ratios. For a standard burst-to-target 
ratio of unity, which is in the neighborhood of service 
conditions, the maximum probability is about 11.0 
times normal, and the probability at the present limit 
of caliber ratios (0.7) is 5.5 times normal. This gain 
in chance of damage is obtained at no increase in am¬ 
munition cost but at a probable actual reduction in 
cost. In other words, the ammunition cost per hit 
will be reduced to less than one-fifth of the present 
cost. By the same token, the same ammunition 
cost can attain damage under much more difficult 
conditions. 


At the present time the life of the gun is reduced to 
about one-third its normal value. By the use of stel¬ 
lite liners (Chapter 19) this can undoubtedly be re¬ 
turned to normal for some guns, or actually increased. 
In other words, almost the entire saving in ammu¬ 
nition cost can be made effective by such design. The 
most important aspect of the situation is, however, 
that a given battery of guns, its directors, auxilliary 
equipment, its whole personnel and the equipment, 
personnel, and personnel and fuel used in transporting 
it to its point of use, are rendered five or more times as 
effective. The material advantages, when viewed from 
this standpoint, are very impressive. 

Limitations on Gains 

While the evaluation of the precise conditions is 
not within the scope of the present discourse, it should 
be pointed out that there is a range for any gun be¬ 
yond which use of a subcaliber projectile will not 
reduce the time of flight. Furthermore, there is a 
shorter range beyond which reduction of caliber does 
not increase the probability of a damaging hit. These 
limitations require thorough analysis in connection 
with any further studies along these lines. 

33 2 4 Gains in Armor Penetration by 
Use of Subcaliber Projectiles 

The de Mar re formula for penetration of armor 
plate by steel projectiles is given by equation (22), 



in which e is the thickness of plate penetrated, d is 
the diameter of the projectile, W is its weight, 0 the 
angle of impact relative to normal, V the striking 
velocity, and a and /3 are constants. For nondeform- 
able projectiles, log a = 6.15 and = 1.43. Hence 
equation (22) reduces to equation (23). 



If the projectile energy is kept constant while its 
diameter is varied, the penetration improves as the 
diameter diminishes. This improvement continues 
until the impact velocity reaches the shatter value. 

When tungsten carbide is used as a projectile ma¬ 
terial, a and j8 both diminish, and the relative pene¬ 
tration is increased. The constants are now, however, 
not independent of 6 and there is a region of 6 wherein 


CONFIDENTIAL 





MILITARY IMPORTANCE OF SUBCALIBER PROJECTILES 


627 


the advantage of tungsten carbide is not sq great. On 
the average, tungsten carbide is in the neighborhood 
of 2.5 to 3.5 times as effective as steel in armor pene¬ 
tration. This gain is due in part to the fact that the 
values of a and 0 are more favorable but also to the 
fact that for a given value of projectile mass, the 
tungsten carbide projectile diameter is only about 
three-fourths that of the steel projectile. 

In view of the higher density, a carbide projectile 
of given diameter has lower velocity than the corre¬ 
sponding steel projectile for given penetration, and 
since it also has a higher shatter velocity, it can strike 
armor plate with much more energy than can a steel 
projectile of the same size. 

The shatter velocity of tungsten carbide is roughly 
three-halves that of steel so a projectile of given 
energy should, on this basis alone, have about two 
and one-half times the penetration of steel. The more 
advantageous values of a and /3 produce the additional 
improvement. 

From equation (23) it is apparent that reduction 
of caliber, while retaining projectile energy, improves 
penetration. However, in an actual gun, the powder 
energy is fixed and, as pointed out in Section 2.2, a 
part of this energy is delivered as kinetic energy of 
the powder gas. In view of this, the energy of the 
projectile diminishes as the caliber is reduced. Fur¬ 
thermore, because of less favorable ballistic condi¬ 
tions, the projectile strikes the plate with a further 
reduction in relative energy and the advantages of 
subcaliber projectiles for armor penetration are there¬ 
by modified. There is, however, a region of advantage 
in which gains in armor-piercing effectiveness are 
actually attained by the use of subcaliber projec¬ 
tiles. 

In order to simplify the present analysis, we neglect 
the mass of skirts and sabots. This is not strictly 
allowable, but the optimum conditions occur in a 
region of values of p for which the error is not serious. 
When this is done, the striking velocity is given by 
equation (24). 

1~ 1 + 0.25( C/Wo) > 

0 Lp 3 + 0.25(C/TFo) J 

/ Kpr fp 3 + 0.25(C/TOT'4 2 (24) 

l 2pVV 0 L 1 + 0.25(C/lfo) J J ' 


Inserting this value for V in equation (23), and divid¬ 
ing by equation (25), 


1 / W 0 V 0 2 cos 2 (A 1 ' 1 - 43 
do 11 \ a J • 


(25) 


equation (26) for the relative penetration is ob¬ 
tained. 


6 T 1 + 0.25(C/Wo) l 1 ' 1 - 43 

e» P Lp 3 + 0.25(C/Wo)J 



K&_ 

2pVF 0 


p 3 + 0.25(C/W o )~Vi i 
. 1 + 0.25( C/Wo) J 

_ AV_ 

2Wo 


4/1.43 


■ (26) 


For values of p near unity, the expression 1 —• A 
may be substituted for p. Upon expansion, and with 
neglect of higher powers of A, equation (26) reduces 
to equation (27). 


e 

eo 


= 1 + A 


- 1 


+ 


1.43(1 + 0.25(C/TFo) 

j_ ( 3 _I_A 

1.43 \4 1 + 0.25(C/ Wo) J 


A'or/2vTo] 

, __K«A (27) 

2VVo) 


It is clear that there is a set of conditions for which 
e/eo is not increased by changing the caliber ratio p. 
These conditions are given by equation (28). 


3 

1.43[1 + 0.25( C/Wo)] 


- 1 = 



K,r_ 

2\/Uo" 


(28) 


When C/Wo is fixed, the range at which no increase 
in penetration results from a reduction of caliber is 
given by equation (29), 


n = 


[2.1/1 + 0.25(C/IFo)] 


1.8 


r 0 = 1.16r 0 


TO. 152 - .12(C/Wo)l 
L 1 + 0.25(C/TFo) J’ 


(29) 


in which ro = 2^Vo/K 0 . For the 120-mm gun, C/Wo 
is approximately 0.45, so the range is 17,000 yd. 

It is thus apparent that gains in armor penetration 
may be accomplished out to ranges which are well 
beyond those at which there would be likely use of a 
gun of this size for the purpose of making direct hits 
against armor. 

The order of magnitude of the gains resulting from 
the use of subcaliber projectiles without change of 
material is indicated in Table 3, which has been pre¬ 
pared for steel projectiles at zero range, for the 
120 -mm gun with standard powder load. 


CONFIDENTIAL 




















628 


PRACTICAL HYPERVELOCITY GUNS 


Table 3. Relative armor penetration from subcaliber 
and standard steel projectiles fired from 120-mm gun at 


zero range. 


p 

e/e 0 

1.0 

1.0 

0.8 

1.16 

0.7 

1.28 

0.6 

1.34 

0.5 

1.40 

0.4 

1.28 


While the gains are not so striking as those result¬ 
ing from the use of tungsten carbide projectiles, they 
are sufficient to be interesting. 

When the diameter of a tungsten carbide projectile 
is 75 per cent that of the standard projectile, its 
weight is substantially the same, so the muzzle ve¬ 
locity does not change. Under these circumstances 
the retardation of the carbide projectile is only of the 
order of half that of the steel projectile. The relative 
penetration is given by equation (30). 


6c _ r i-(K 0 r/4VU 0 > 

eo ~ Ll - (K 0 r/ 2Fo)_r 


(30) 


Thus we have the remarkable condition that not only 
is the initial penetration of the tungsten carbide 
greater, but it improves materially with range. In the 
case of the 120-mm gun, with an equal weight tung¬ 
sten carbide projectile the relative penetration for 
different ranges is given in Table 4. The absolute limit 


Table 4. Relative armor penetration from tungsten 
carbide-cored subcaliber projectiles and standard pro¬ 
jectiles fired from 120-mm gun. 


r (yd) 

e/e 0 

0 

1.4 

5,000 

2.0 

10,000 

3.0 

15,000 

5.7 

20,000 

13.0 


of range for this gun is about 30,000 yd with the steel 
projectile, and double that with the carbide projectile. 
The normal ballistic tables extend to about 17,000 
yd. This is perhaps as striking an indication of the 
advantage of the combination of subcaliber and tung¬ 
sten carbide in projectile design as can be given, for 
it combines the effects of increased normal penetra¬ 
tion and improved ballistics. 

As the weight of the tungsten carbide projectile is 
further reduced, the initial penetration increases and 
for a while there is a continued increase all down the 


line. Ultimately, a point is reached where the ratio 
of tungsten carbide velocity to steel velocity is con¬ 
stant throughout the range of travel, and under this 
condition the penetration ratio is constant. This oc¬ 
curs, however, at a value of p less than 0.2 which is 
well below the practical limit. 

In some respects absolute values of penetration are 
more significant than relative values. From this stand¬ 
point it is worth noting that, in the case of the 
120 -mm gun, the equal-weight subcaliber projectile 
has, at a range of about 2,800 yd, the same penetra¬ 
tion as the steel projectile at zero yd. At 5,000 yd the 
tungsten carbide penetration is 85 per cent that for 
steel at zero range, while at 10,000 yd it is 72 per cent, 
and at 15,000vd 53per cent. At 15,000 yd the penetra¬ 
tion for steel is only 9 per cent of that at zero range. 


33.3 WHAT OF THE FUTURE? 

33 3 1 Continuation of Fundamental 
Research 

Although Division 1, by its intensive effort during 
a period of five years, made gratifying progress to¬ 
ward the understanding of erosion and the practical 
solution of the hypervelocity problem, it cannot be 
claimed that the problem is completely solved. The 
investigations uncovered a number of new problems 
interesting in themselves and of fundamental impor¬ 
tance in connection with further improvements in 
gun design. Nearly all parts of the broad program of 
Division 1 could be carried further profitably by the 
Army or Navy or by an independent organization 
operating in conjunction with the Armed Services. In 
the future, great emphasis should continue to be 
placed upon basic research. The accomplishments of 
Division 1 have resulted mainly from an understand¬ 
ing of the fundamentals and not from a mere “polish¬ 
ing up” of previously existing devices. Developing 
and testing are of course essential, but substantial 
improvements depend on the new ideas that evolve 
as the result of research. Some of the fields on which 
it appears emphasis should be placed in future in¬ 
vestigations are mentioned below. 

33 3 2 Interior Ballistics 

Although important advances have been made 
recently in this field, additional information is needed. 
In particular, an effort should be made to secure 


CONFIDENTIAL 










WHAT OF THE FUTURE? 


629 


simplified relations between the various factors in¬ 
volved. Furthermore, hypervelocity gun design would 
be facilitated by improvement of ignition and insur¬ 
ance of uniform ballistics even at very high densities 
of loading and correspondingly high pressures. 

33 3 3 Heating of Guns 

Much additional work needs to be done before we 
have a complete knowledge of the temperatures in 
various parts of a gun barrel under varying conditions 
of firing. Further investigations should include both 
experimental determinations and mathematical anal¬ 
yses. It is essential to know more completely the 
temperatures at the bore surface of guns of various 
sizes under various conditions of firing, in order to 
deal intelligently with all aspects of gun erosion; and 
it is desirable to have full knowledge of the tempera¬ 
tures at points in the wall in order to predict the 
behavior of new designs of gun barrels. Such informa¬ 
tion is important also in connection with the deter¬ 
mination of the value of cooling devices. 

33 3 4 Erosion 

Although we now have a fairly satisfactory picture 
of gun erosion, further studies would be worth while. 
Laboratory apparatus that would separate properly 
the various phenomena involved is desirable. For 
example, an improvement in the electron bombard¬ 
ment technique (Section 11.3.2) for producing ther¬ 
mal effects that may be completely free from chemical 
ones would be useful. It is interesting to note that 
although many kinds of laboratory tests relating to 
erosion have been devised, we still do not have the 
means of duplicating at will in a simple laboratory 
apparatus the erosion that takes place in an actual 
gun. So long as it continues to be necessary to test 
erosion-resistant materials, there will be a need for 
laboratory apparatus by which we may avoid part or 
all of the cumbersome and expensive testing that now 
must be carried out with guns, which for the larger 
sizes involves very considerable cost in money and 
time. 

33 3 5 Bore Friction 

Much effort would be justified in an attempt to 
secure more complete information on bore friction in 
various types and sizes of guns. This investigation 
might well be carried forward as a part of a long-range 


research program of investigations in interior ballis¬ 
tics, using all possible experimental and theoretical 
approaches. Rapid and convenient techniques for 
measuring bore friction would facilitate the develop¬ 
ment of improved rifling and banding’ and in general 
of methods for reducing friction, which is the direct 
or indirect source of various difficulties with hyper¬ 
velocity guns. Further data would be useful in de¬ 
veloping methods for reducing friction by improved 
band design or otherwise. 

33 3 6 Liners and Coatings 

Recent results on the selection and application of 
liners and coatings of erosion-resistant materials are 
encouraging and justify further studies. Adequate 
firing tests should be made of materials such as hard¬ 
ened molybdenum, chromium-base alloys, and elec¬ 
troplated duplex coatings or alloy coatings especially 
of cobalt and tungsten. Although the low fusion point 
of stellite appears to limit its utility to small guns and 
moderate muzzle velocities, some further tests would 
be worth while. The theoretical possibilities of a hot- 
hard alloy such as stellite coated with a refractory 
material like molybdenum or chromium are so attrac¬ 
tive that this particular development should be con¬ 
tinued until it is determined whether or not this 
combination will afford a universal solution of the 
hypervelocity problem, by preventing both erosion 
and flattening of the lands under the most severe 
conditions. In this group of projects may also be 
placed further studies on the Fisa protector (Chapter 
32), the advantages of which, either alone or in com¬ 
bination with other features, have not been fully 
evaluated. 

33 3 7 Improved Propellants 

It appears quite obvious that it would be well 
worth while to attempt to develop propellants that 
either by low flame temperature or favorable chem¬ 
ical composition will be less erosive (in proportion to 
energy) than existing propellants. Furthermore, ade¬ 
quate tests should be made on mixtures of single-base 
and double-base propellants having an approximately 
“neutral” chemical action, that is, ones in which the 
carburization of the bore is reduced without excessive 
increase of flame temperature (Section 15.6). There 
remains a possibility that by minimizing chemical 
action and the production of low-melting compounds 
on the bore surface, the erosion-resistance of gun steel 


CONFIDENTIAL 



630 


PRACTICAL HYPERVELOCITY GUNS 


could be significantly improved. Further investiga¬ 
tions using an adiabatic compression apparatus (Sec¬ 
tion 11.3.1) might provide a useful clue as to desired 
compositions for nonerosive propellants. 

33 3 8 Special Projectiles 

Sabot-projectiles (Chapter 29) have been devel¬ 
oped that have satisfactory accuracy and meet the 
requirement of firing through a muzzle brake; but 
further improvements are desirable, and additional 
studies should be made so as to determine the best 
types for various uses. The utility of tapered-bore 
guns and skirted projectiles (Chapter 30) is worthy 
of further evaluation as emphasized in Section 33.2. 
The pre-engraved projectile in combination with a 
chromium-plated barrel (Chapter 31), which tests in 
a small-caliber gun indicate is a remarkably effective 
method for reducing gun erosion and thus making 
hypervelocity practical, should be studied further in 
order to ascertain what conditions of plating are 
necessary for large guns and to find the most suitable 
projectile designs. In this connection, it is vital to 
develop simple and convenient loading and indexing 
mechanisms (Section 28.4) in order to utilize fully the 
advantages of pre-engraved projectiles for various 
types of guns. Another item that would justify criti¬ 
cal evaluation is the “booster” projectile, especially 
the type which near the end of its flight receives an 
additional impulse from a rocket-like mechanism. e 


e A similar suggestion is given by Division 8 in its section of 
the History of OSRD. 


33.3.9 High-Strength Steels for Gun 

Barrels 

The use of much higher pressures for hypervelocity 
guns is an interesting possibility. This would simplify 
the design of such guns, but would serve to emphasize 
the desirability of making gun tubes of steel having a 
strength much higher than that of ordinary gun steel. 
There is good reason to believe that the avoidance of 
damage to the bore surface through erosion would 
greatly increase the probability of favorable behavior 
with steels of high strength. 

33.3.10 "Economics” of Gun Fire 

From a broad point of view, the overall advantage 
of hyper velocity is to be judged by performance in 
relation to weight, cost, and mobility. It is vital, 
therefore, to make a careful determination of the 
optimum conditions for hypervelocity guns so as to 
find the combination of factors that involves the 
minimum consumption of labor and critical material 
per unit of damage. As a result of the new knowledge 
accumulated in the last five years, there can remain 
little doubt that the guns of the future will operate 
at velocities higher than in the past. Hypervelocity 
is advantageous and it is practicable. A practical 
limitation on muzzle velocity is now imposed not by 
physics, but rather by economics. Whether the limit 
beyond winch it will be unprofitable to increase the 
velocity is generally at 4,000 or 4,500 fps or much 
higher is to be determined by future assessment and 
balancing of all the relevant factors. 


CONFIDENTIAL 




GLOSSARY 


Accuracy Life. The number of rounds that can be fired 
from a gun before it becomes inaccurate because of erosion. 
The criterion for accuracy-life varies from gun to gun. Cf. 
velocity life. (See Section 23.1.4.) 

Aircraft Barrel, caliber .50. A 36-in. caliber .50 machine 
gun barrel made according to Ordnance Department Draw¬ 
ing D35348A or D28272. 

Altered Layer. A thin layer of steel at the bore surface of a 
gun, the chemical and physical nature of which differs from 
that of the underlying steel because of changes that have 
taken place during firing. (See Section 12.1.1.) 

Austenite. A solid solution of carbon in the high-tempera¬ 
ture form of iron, called gamma-iron. Cf. ferrite. (See Section 
12 . 2 . 1 .) 

Ballistic Level. The charge and pressure at which a gun 
is fired. 

Body Engraving. A mark or marks on the bourrelet or body 
of a projectile caused by contact with the lands of the gun. 
(See Sections 6.1.2, 10.4.10.) 

Bore Profile. A graph in which is plotted either the bore 
diameter or the change in bore diameter of a gun as a func¬ 
tion of the position along the axis of the bore. (See Section 
10 . 2 . 1 .) 

Cannelure. A circumferential groove in a projectile. 

Cementite. Iron carbide, one of the constituents of steel, 
which is sometimes formed on the bore surface of a gun 
during firing. (See Section 12.5.2.) 

Clock Position. A position on the circumference of the bore 
of a gun designated by analogy with the position of the hour 
hand on the face of a clock. 

Cook-Off. The premature explosion of the powder charge in 
a gun chamber or of the high-explosive filler in a shell or fuze 
caused by the heat of the gun w'hen the round is left in the 
chamber. (See Section 5.6.2.) 

Co volume. A correction factor in the equation of state of 
the powder gases w r hich takes into account the space oc¬ 
cupied by the gas molecules themselves. (See Section 2.4.2.) 

Cyclonite. A high explosive that has been used experi¬ 
mentally as the active ingredient in a propellant by dis¬ 
persing the crystals of the compound in a nitrocellulose 
matrix. 

DPH. “Diamond pyramid hardness,” which is measured by 
means of a Vickers hardness tester. 

Double-base Powder. A powder containing nitroglycerine 
and nitrocellulose as its active ingredients. 

Duplex Plate. An electroplate consisting of two or more 
separate electroplates deposited one on top of the other. 
(See Section 20.2.4.) 


ESR. “Equivalent Service Rounds,” that is, the number of 
full-charge rounds that is considered equivalent to a larger 
number of rounds fired at a reduced charge. (See Section 
10.3.3.) 

Ferrite. A solid solution of not more than 0.06% carbon in 
the low'-temperature form of iron, called alpha-iron. Cf. 
austenite. (See Section 12.2.1.) 

Fisa Protector. A device for protecting the bore of a gun 
from erosion. (See Chapter 32.) 

Heavy Barrel, Caliber .50. A 45-in. caliber .50 machine 
gun barrel made according to one of the following Ordnance 
Department drawings: D28253-11, D28253A, D28253A3, 
or D28269-8X. 

HC Plate. “High contraction plate,” a type of chromium 
plate. (See Section 20.2.2.) 

Hypervelocity. A projectile velocity greater than those 
commonly used, arbitrarily taken by Division 1 as one in 
excess of 3,500 fps. (See Section 1.1.3.) 

LC Plate. “Low contraction plate,” type of chromium 
plate. (See Section 20.2.2.) 

Microflash. A high-intensity light source of very short 
duration used for photographing moving objects, manufac¬ 
tured by the General Radio Company after a design devel¬ 
oped by H. E. Edgerton. 

NC. Nitrocellulose. 

NG. Nitroglycerine. 

Nital. A solution of nitric acid in ethyl alcohol used to etch 
steel specimens prior to metallographic examination. 

Obturation. Sealing of a gun bore against escape of the 
pow r der gases, either through the breech or past the projec¬ 
tile. 

Parco-Lubrizing. A trade-marked process (Parker Rust 
Proof Company, Detroit, Michigan) for applying to steel a 
w ear-resistant coating that consists chiefly of an admixture 
of iron and manganese phosphates. It was found advan¬ 
tageous to use it on pre-engraved projectiles. 

Pebbling. The roughness that appears on the bore surface 
of a gun that has been liquefied by the powder gases. (See 
Section 13.5.3.) 

RDX. A code name for cyclonite, the active ingredient of 
some propellants. 

RPM. Rounds per minute. 

Velocity Life. The number of rounds that can be fired from 
a gun before its muzzle velocity decreases a predetermined 
amount because of erosion. A value of 200 fps was adopted 
for this velocity drop for the caliber .50 aircraft barrel 
during World War II. Cf. accuracy life. (See Section 23.1.4.) 


CONFIDENTIAL 


631 









BIBLIOGRAPHY 


Numbers such as Div. 1-540-MI indicate that the document listed has been microfilmed and that its title appears in the micro¬ 
film index printed in a separate volume. For access to the index volume and to the microfilm, consult the Army or Navy agency 
listed on the reverse of the half-title page. 

Some abbreviations are used in identifying the source of reports. They include the following: 

CIW Carnegie Institution of Washington 

NBS National Bureau of Standards 

MIT Massachusetts Institute of Technology 

BTL Bell Telephone Laboratories, Inc. 

BRL-APG Ballistic Research Laboratory, Aberdeen Proving Ground 

WA Watertown Arsenal 

OCO Office of the Chief of Ordnance 

AC Advisory Council on Scientific and Technical Development 

ADD Armament Design Department 

ARD Armament Research Department t 

RD Research Department 

The number prefixed by the letter “A” following the OSRD number is the number assigned by the Technical Reports Section, 
NDRC (formerly in the office of Division A, NDRC) at the time the report was edited and prepared for duplication by that 
office. A number of these reports had not been duplicated by the time the office of Division A was closed. Manuscript copies 
were supplied for microfilming. 


FORMAL REPORTS ISSUED BY DIVISION 1, NDRC 


1. Methods for Detecting Defects in the Base of a Shell Forg¬ 

ing, R. W. Goranson, OSRD 373, Report A-31, CIW, 
Feb. 6, 1942. Div. 1-540-MI 

2. A Brief History of Tapered Bore Guns, J. S. Burlew, 

OSRD 515 Preliminary Report A-43, Geophysical Lab., 
CIW, Apr. 17, 1942. Div. 1-330-MI 

3. The Measurement of Large Transient Stresses, R. W. Gor¬ 
anson, W. Garten, Jr., and J. A. Crocker, OSRD 498, 
Progress Report A-45, Geophysical Lab., CIW, Apr. 10, 

1942. Div. 1-210.1-MI 

4. A Method of Investigating the Deformation of Deformable 
Projectiles, H. L. Whittemore and L. R. Sweetman, 
OSRD 631, Memorandum A-36M, NBS, July 16, 1942. 

Div. 1-540-M2 

5. The Piezoelectric Projectile Accelerometer and a Bore-Fric¬ 
tion Gage, N. M. Smith, Jr., OSRD 808, Progress Report 
A-59, Geophysical Lab., CIW, Aug. 17, 1942. 

Div. 1-210.31-MI 

6. Simple Calculation of Thermochemical Properties for Use 
in Ballistics, Addenda to NDRC Report A-101, J. O. 
Hirschfelder and J. H. Sherman, OSRD 1300, Memo¬ 
randa A-67M to A-70M, Geophysical Lab., CIW March 

1943. Div. 1-210.2-M4 

7. A Vertical Step-sweep Circuit for the Cathode-Ray Oscillo¬ 
graph, N. M. Smith, Jr., OSRD 2081, Memorandum 
A-77M, Geophysical Lab., CIW, November 1943. 

Div. 1-620-MI 

8. The Preparation of Chromium by the Thermal Decompo¬ 
sition of Chromium Iodide, D. R. Mosher, OSRD 2082, 
Memorandum A-78M, Westinghouse Electric and Manu¬ 
facturing Co., Inc., November 1943. Div. 1-420.33-MI 

9. Thermodynamic Properties of British Flashless and Cordite 
MD Powders, F. T. McClure, D. W. Osborne, and J. O. 


Hirschfelder, OSRD 817, Progress Report A-82, Geo¬ 
physical Lab., CIW, Aug. 24, 1942. Div. 1-210.2-MI 

10. A Theorem on Radial Deformation in Thick Tubes, K. F. 
Herzfeld, OSRD 3209, Memorandum A-83M, Catholic 
University of America, January 1944. Div. 1-320-MI 

11. A New Thread Collocating Gage, F. E. Blake and D. F. 
Ringie, OSRD 3287, Memorandum A-85M, Jones & 
Lamson Machine Co., February 1944. Div. 1-620-M2 

12. Heat Conduction, Gas Flow, and Heat Transfer in Guns, 

J. O. Hirschfelder, W. Garten, Jr., and O. Hougen, OSRD 
863, Progress Report A-87, Geophysical Lab., CIW, 
Sept. 1, 1942. Div. 1-310-MI 

13. On Longitudinal Stresses in Guns, K. F. Herzfeld, OSRD 

3421, Memorandum A-87M, Catholic University of 
America, March 1944. Div. 1-320-M2 

14. Stability of Subcaliber Projectiles, C. L. Critchfield, OSRD 

870, Progress Report A-88, Geophysical Lab., CIW, 
Sept. 8, 1942. Div. 1-510-MI 

15. The Erosion of Guns, Part I, J. S. Burlew, OSRD 882, 

Progress Report A-90, Geophysical Lab., CIW, Sept. 15, 
1942. Div. 1-400-M2 

16. An Improved Approximation Formula for Stresses in Cyl¬ 
inders, K. F. Herzfeld, OSRD 3465, Memorandum A- 
90M, Catholic University of America, April 1944. 

Div. 1-320-M3 

17. A Comparison of Ballistic Systems, J. O. Hirschfelder, 
R. B. Kershner, C. F. Curtiss, J. Sherman, R. E. John¬ 
son, N. L. Johnson, J. W. Wrench, Jr., E. Harmon, and 
M. G. Aldrich, OSRD 3556, Memorandum A-91M, 
Geophysical Lab., CIW, April 1944. Div. 1-210.1-M6 

18. Interaction of Carbon Monoxide and Iron, J. C. W. Frazer 
and F. H. Horn, OSRD 873, Progress Report A-92, 
Johns Hopkins University, Sept. 12, 1942. 

Div. 1-410.1-MI 


CONFIDENTIAL 


633 


634 


BIBLIOGRAPHY 


19. Apparatus far Collecting Solid Particles Discharged 
from a Rifle. A Simple Optical Sighting Device, J. L. 
England, OSRD 3753, Memoranda A-93M & A-94M, 
Geophysical Lab., CIW, May 1944. Div. 1-420.23-M2 

20. Notes on the Potentiometric Titration of Iron, E. Jensen, 

OSRD 3869, Memorandum A-95M, Geophysical Lab., 
CIW, July 1944. Div. 1-650-MI 

21. The Erosion of Guns, Part II. J. S. Burlew, OSRD 982, 

Progress Report A-91, Geophysical Lab., CIW, Nov. 4, 
1942. Div. 1-400-M3 

22. Anodic Polishing for the Removal of Very Thin Layers 

from Steel Surfaces, W. D. Urry and E. Jensen, OSRD 
3968, Memorandum A-96M, Geophysical Lab., CIW, 
July 1944. Div. 1-650-M2 

23. Sintple Calculation of Thermochemical Properties for Use 

in Ballistics, J. O. Hirschfelder and J. Sherman, OSRD 
935, Progress Report A-101, Geophysical Lab., CIW, 
Oct. 14, 1942. Div. 1-210.2-M2 

24. a Thermodynamic Properties of Propellant Gases, J. O. 

Hirschfelder, F. T. McClure, C. F. Curtiss, and D. W. 
Osborne, OSRD 1087, Progress Report A-116, Geophysi¬ 
cal Lab., CIW, Nov. 25, 1942. Div. 1-210.2-M3 

25. The Drag Coefficient for a Cone Moving with High Ve¬ 
locity, W. Karush and C. L. Critchfield, OSRD 1104, 
Report No. A-126, Geophysical Lab., CIW, Dec. 21,1942. 

Div. 1-220.1-MI 

26. Interior Ballistics [Part I], J. O. Hirschfelder, R. B. 
Kershner, and C. F. Curtiss, OSRD 1236, Report No. 
A-142, Geophysical Lab., CIW, February 1943. 

Div. 1-210.1-M2 

27. Metals Tested as Erosion Vent Plugs, O. H. Loeffler, G. 
Phair, and H. S. Jerabek, OSRD 1249, Report No. A-148, 
Geophysical Lab., CIW, Feb. 19, 1943. 

Div. 1-420.21-MI 

28. A Physico-Chemical Study of Gun Erosion, E. Posnjak, 

OSRD 1311, Report No. A-161, Geophysical Lab., 
CIW, March 1943. Div. 1-410.1-M2 

29. Velocity and Pressure during Free Run-Up of a Projectile, 

C. F. Curtiss, OSRD 1430, Report No. A-179, Geophysi¬ 
cal Lab., CIW, April 1943. Div. 1-210.1-M3 

30. Interior Ballistics [Part II], J. O. Hirschfelder, R. B. 
Kershner, C. F. Curtiss and J. Sherman, OSRD 1435, 
Report No. A-180, Geophysical Lab., CIW, April 1943. 

Div. 1-210.1-M2 

31. A Study by Means of Electron and X-Ray Diffraction of 
the Alteration of Steel by Hot Powder Gases, L. H. Germer, 
J. J. Lander, G. Tunell, and P. N. Metzelaar, OSRD 
1659, Report No. A-199, Bell Telephone Laboratories, 
Inc., and Geophysical Lab., CIW, July 1943. 

Div. 1-410.1-M4 

32. The Temperature of the Bore Surface of Guns, G. S. Ful¬ 

cher (editor), OSRD 1666, Report No. A-201, Geophysi¬ 
cal Lab., CIW, July 1943. Div. 1-310-M2 


a A previous edition of this report had been designated as 
Report A-48 (OSRD 547). 


33. Interior Ballistics [Part III], J. O. Hirschfelder, R. B. 
Kershner, and J. H. Sherman, OSRD 1677, Report No. 
A-204, Geophysical Lab., CIW, July 1943. 

Div. 1-210.1-M2 

34. Exact Theory of the Stress Distribution in a Shell Due to En¬ 

graving, H. A. Jordan, K. F. Herzfeld, and V. O. Mc- 
Brien, OSRD 1714, Report No. A-207, Catholic Uni¬ 
versity of America, July 1943. Div. 1-520-MI 

35. Interior Ballistics [Part IV], C. F. Curtiss and R. E. 

Johnson, OSRD 1740, Report No. A-208, Geophysical 
Lab., CIW, August 1943. Div. 1-210.1-M2 

36. Interior Ballistics of Recoilless Guns, J. O. Hirschfelder, 

R. B. Kershner, C. F. Curtiss, and R. E. Johnson, OSRD 
1801, Report No. A-215, Geophysical Lab., CIW, Sep¬ 
tember 1943. Div. 1-210.1-M4 

37. Interior Ballistics [Part V], The Performance of High- 
Velocity Guns, J. O. Hirschfelder, R. B. Kershner, C. F. 
Curtiss, and R. E. Johnson, OSRD 1916, Report No. 
A-222, Geophysical Lab., CIW, October 1943. 

Div. 1-210.1-M5 

38. Metals Tested for Resistance to Cavitation Erosion, G. E. 
Ziegler and L. E. Line, OSRD 1917, Report No. A-223, 
Armour Research Foundation, October 1943. 

Div. 1-420.23-MI 

39. Report on Firing of First Eleven Rounds in 3-In. Gun at 
David W. Taylor Model Basin during May, June, July 
1943, OSRD 2019, Report No. A-229, NBS and Geophys¬ 
ical Lab., CIW, November 1943. Div. 1-210.3-MI 

40. Carbon and Nitrogen in Gun Erosion, J. F. Schairer and 

E. G. Zies, OSRD 2042, Report No. A-230, Geophysical 
Lab., CIW, November 1943. Div. 1-410.1-M5 

41. Development of Subcaliber Projectiles for the Hispano- 

Suiza Gun, C. L. Critchfield and J. McG. Millar, OSRD 
2067, Report No. A-233, Geophysical Lab., CIW, No¬ 
vember 1943. Div. 1-510-M2 

42. Sabot-Projectiles for Cannon, W. D. Crozier, H. F. Dun¬ 

lap, C. E. Hablutzel, L. Lepaz, and D. T. MacRoberts, 
OSRD 3010, Report No. A-234, University of New Mex¬ 
ico, December 1943. Div. 1-510.1-MI 

43. A Method of Obtaining the State and Composition of the 

Powder Gas at Shot Ejection with Tables for Pyro Powder, 
R. E. Johnson, OSRD 3239, Report No. A-248, Geophys¬ 
ical Lab., CIW, February 1944. Div. 1-210.2-M5 

44. An Experimental Study of Powder Gas Radiation and 

Temperature, F. C. Kracek and W. S. Benedict, OSRD 
3291, Report No. A-252, Geophysical Lab., CIW, Feb¬ 
ruary 1944. Div. 1-210.2-M6 

45. On the Heating of Rotating Bands, C. L. Critchfield, 

OSRD 3329, Report No. A-256, Geophysical Lab., 
CIW, February 1944. Div. 1-530-MI 

46. Measurement of Various Ballistic Quantities on a Pro¬ 
jectile Moving in the Bore of a Gun, N. M. Smith, Jr., 
and J. A. Crocker, OSRD 3376, Report No. A-259, Geo¬ 
physical Lab., CIW, March 1944. Div. 1-210.3-M2 

47. The Pressure on a Cone Moving with Small Yaw at High 
Velocity, W. Karush and C. L. Critchfield, OSRD 3397 


CONFIDENTIAL 





BIBLIOGRAPHY 


635 


Report No. A-260, Geophysical Lab., CIW, March 1944. 

Div. 1-220.1-M2 

48. Thermal Effects of Propellent Gases in Erosion Vents and 

in Guns, L. W. Nordheim, H. Soodak, and G. Nordheim, 
OSRD 3447, Report Nq. A-262, Duke University, May 
1944. Div. 1-310-M3 

49. Investigation of Gun Erosion at the Geophysical Laboratory, 
Vol. 1, July 1941 to July 1943, OSRD 3448, Report No. 
A-263, Geophysical Lab., CIW, March 1944. 

Div. 1-400-MI 

50. Investigation of Gun Erosion at the Geophysical Labora¬ 
tory, Vol. II, July 1943 to December 1943, OSRD 3449, 
Report No. A-264, Geophysical Lab., CIW, April 1944. 

Div. 1-400-MI 

51. A Method for Testing Resistance of Metals to Surface 

Cracking under Conditions Similar to Those Obtaining in 
Guns, E. Ingerson, OSRD 3628, Report No. A-271, Geo¬ 
physical Lab., CIW, May 1944. Div. 1-410.3-MI 

52. Preliminary Report on Molybdenum as a Material for an 
Erosion-Resistant Gun Liner, P. H. Brace and J. W. Mar- 
den, OSRD 3700, Report No. A-273, Westinghouse 
Electric and Manufacturing Co., Inc., May 1944. 

Div. 1-420.32-MI 

53. The Effect of Sulfur and Other Components of Black Pow¬ 

der on the Erosion of Gun Steel, W. D. Urry and E. Inger¬ 
son, OSRD 3811, Report No. A-276, Geophysical Lab., 
CIW, June 1944. Div. 1-410.1-M6 

54. Molding Sabots for Projectiles, OSRD 3832, Report No. 
A-278, Arthur D. Little, Inc., June 1944. 

Div. 1-510.1-M2 

55. Interior Ballistics [Part VI], Pressure Travel Curves, with 
an Addendum to Interior Ballistics, /., R. E. Johnson, 
C. F. Curtiss, and R. B. Kershner, OSRD 3855 Report 
No. A-279, Geophysical Lab., CIW, June 1944. 

Div. 1-210.1-M7 

56. Deduction of Practical Formulas for the Stress in the Mantle 
of a Shell Due to Band Pressure and Powder-Gas Pressure, 
K. F. Herzfeld and V. Griffing, OSRD 3868, Report No. 
A-281, Catholic University of America, June 1944. 

Div. 1-530-M2 

57. Trajectory Determination by Tracer Photography, W. D. 

Crozier, OSRD 3890, Report No. A-283, University of 
New Mexico July 1944. Div. 1-220.2-MI 

58. Formulae for Strains in a Thick-Walled Tube near the 

Projectile, C. Snow, OSRD 4320, Report No. A-298, 
NBS, November 1944. Div. 1-320-M4 

59. Investigation of Gun Erosion at the Geophysical Labora¬ 

tory, Vol. Ill, January 1944 to July 1944, OSRD 4345, 
Report No. A-300, Geophysical Lab., CIW, November 
1944. Div. 1-400-MI 

60. Chemical Thermodynamics of Gun Erosion, C. F. Curtiss 
and N. L. Johnson, OSRD 4363, Report No. A-301, Geo¬ 
physical Lab.,* CIW, November 1944. Div. 1-410.1-M8 

61. The State of Equilibrium among the Carbon Atoms of a 

Propellant Gas, W. D. Urry and P. J. Hannan, OSRD 
4461, Report No. A-303, Geophysical Lab., CIW, Decem¬ 
ber 1944. Div. 1-210.2-M7 


62. Vent Plug Erosion by the CO-CO 2 Gas System, J. C. W. 
Frazer, F. H. Horn, and R. C. Evans, OSRD 6327, Re¬ 
port No. A-310, Johns Hopkins University, Oct. 31, 1944. 

Div. 1-420.21-M2 

63. Iron-Carbonyl Formation as a Mechanism Contributing to 

Gun Erosion, J. C. W. Frazer, F. H. Horn, and R. C. 
Evans, OSRD 6328, Final Report No. A-311, Johns Hop¬ 
kins University, Oct. 31, 1944. Div. 1-410.1-M7 

64. Static Band Pressures in 37-MM Projectiles [Part I], F. A. 

Biberstein, Jr., R. Brown, V. Griffing, K. F. Herzfeld, 
J. M. Krafft, and W. T. Whelan, OSRD 4550, Report 
No. A-312, Catholic University of America, January 
1945. Div. 1-530-M3 

65. Second Report on Firings in 3-In. Gun at David W. Taylor 

Model Basin, October 1943 to May 1944, OSRD 4986, 
Report No. A-323, NBS and Geophysical Lab., CIW, 
Mar. 25, 1945. Div. 1-210.3-M3 

66. Interior Ballistics [Part VII], Numerical Methods of 

Solution of the Ordinary Problems of Interior Ballistics, 
R. E. Johnson, N. L. Johnson, and J. W. Wrench, Jr., 
OSRD 6231, Report No. A-348, Geophysical Lab., 
CIW, Oct. 1, 1945. Div. 1-210.1-M9 

67. The Erosion of Guns at the Muzzle, L. E. Line, Jr., OSRD 
6322, Report No. A-357, Aug. 23, 1945. Div. 1-430-M2 

68. The Aerodynamics of a Slightly Yawing Supersonic Cone, 

A. H. Stone, OSRD 6306, Report No. A-358, Geophysi¬ 
cal Lab., CIW, July 10, 1945. Div. 1-220.1-M3 

69. Interior Ballistics, A Consolidation and Revision of Pre¬ 

vious Reports, Interior Ballistics [Parts I to VII, Inclu¬ 
sive], C. F. Curtiss and J. W. Wrench, Jr., OSRD 6468, 
Report No. A-397, Geophysical Lab., CIW, July 15, 
1945. Div. 1-210.1-M8 

70. The Penetration of Nitrogen into Steel Rifle Barrels as 

Measured by a Tracer Method, G. L. Davis, OSRD 6469, 
Report No. A-398, Geophysical Lab., CIW, Nov. 30, 
1945. Div. 1-410.1-M10 

71. Measurement of Heat Input to the Bore Surface of Caliber 
.50 Gun Barrels, E. L. Armi, J. L. Johnson, R. C. Mach- 
ler, and N. E. Polster, OSRD 6470, Report No. A-399, 
Leeds & Northrup Co., July 23, 1945. Div. 1-310-M4 

72. Measurement of Frictional Heat Input in Gun Barrels and 

of Frictional Bullet-Bore Interface Temperatures, E. L. 
Armi, J. L. Johnson, R. C. Machler, and N. E. Polster, 
OSRD 6471, Report No. A-400, Leeds & Northrup Co., 
Sept. 27, 1945. Div. 1-310-M5 

73. The Synthesis of Chromium Hexacarbonyl, B. B. Owen, 

OSRD 6472, Report No. A-401, Yale University, Sept. 
17, 1945. Div. 1-630-MI 

74. Pyrolytic Plating of Chromium from the Vapor of Chro¬ 
mium Hexacarbonyl, B. B. Owen, OSRD 6473, Report- 
No. A-402, Yale University, Oct. 8,1945. Div. 1-630-MI 

75. The Results of Erosion Vent-Plug Tests Particularly un¬ 
der Conditions of Decreased Severity and Their Applica¬ 
tion to the Erosion of Guns, H. S. Jerabek, G. Phair, D. 
Enagonio, and C. A. MacQuaid, OSRD 6474, Report 
No. A-403, Geophysical Lab., CIW, Dec. 5, 1945. 

Div. 1-420.21-M3 


CONFIDENTIAL 



636 


BIBLIOGRAPHY 


76. The Behavior of Gun Liners and Coatings Tested under 
Conditions of Hypervelocity, N. H. Smith, OSRD 6475, 
Report No. A-404, Franklin Institute, Oct. 2, 1945. 

Div. 1-340-MI 

77. Metallographic Examination of Gun Liners and Coatings 

Tested under Conditions of Hypervelocity, J. N. Hobstet- 
ter, OSRD 6476, Report No. A-405, Harvard Univer¬ 
sity, Oct. 30, 1945. Div. 1-340-M2 

78. Erosion Tests of Materials in the Form of Short Liners in 

a Caliber .30 Machine-Gun Barrel, J. Wulff, OSRD 6477, 
Report No. A-406, Johnson Automatics, Inc., Apr. 12, 

1944. Div. 1-420.3-MI 

79. Search for Erosion-Resistant Materials for Guns by Firing 

Particles of Metal and Alloys into Vacuum to Determine 
Their Structural and Chemical Behavior, E. Posjnak, 
OSRD 6478, Report No. A-407, Geophysical Lab., CIW, 
Dec. 4, 1945. Div. 1-420.23-M3 

80. Gun-Barrel Liners — Materials, Insertion, and Testing, 
F. D. Cotterman, N. A. Ziegler, and J. P. Magos, OSRD 
6479, Report No. A-408, Crane Company, Jan. 16, 1946. 

Div. 1-420.3-M2 

81. The Testing of Erosion-Resistant Materials and the Devel¬ 

opment of Improved Machine-Gun Barrels, E. F. Osborn, 
OSRD 6480, Report No. A-409, Geophysical Lab., 
CIW, Nov. 29, 1945. Div. 1-420.2-MI 

82. Progressive Centrifugal Inmelting for the Preparation of 

Alloy Tubes, P. H. Brace, OSRD 6481, Report No. 
A-410, Westinghouse Electric and Manufacturing Co., 
Inc., Apr. 8, 1946. Div. 1-640-MI 

83. Chromium and Chromium-Base Alloys as Materials for 
Gun Liners, P. H. Brace, J. F. Schairer, and N. A. Zieg¬ 
ler, OSRD 6482, Report No. A-411, Westinghouse 
Electric and Manufacturing Co., Inc., Jan. 5, 1946. 

Div. 1-420.33-M2 

84. Experimental Electroplating of Gun Barrels, W. Blum, 

A. Brenner, and V. A. Lamb, OSRD 6483, Report No. 
A-412, NBS, Dec. 21, 1945. Div. 1-420.4-M3 

85. An Illustrated Study of the Effects of Firing on Chromium- 
Plated Bores of Caliber .50 Machine Guns, H. E. Merwin 
and M. Sullivan, OSRD 6484, Report No. A-413, Geo¬ 
physical Lab., CIW, December 1945. Div. 1-420.4-M2 

86. Symposium on Chromium-Plating, Washington, D. C., 

May 14 , 1943, Division One, NDRC, OSRD 6485, Re¬ 
port No. A-414. Div. 1-420.4-MI 

87. Development of Chromium-Base Hot-Hard Alloys as Gun 

Liner Materials, R. M. Parke and F. P. Bens, OSRD 
6486, Report No. A-415, Climax Molybdenum Co., 
Jan. 21, 1946. Div. 1-420.33-M3 

88. Stellite No. 21 as a Material for Gun Liners—Metallurgy 

and Properties, W. A. Wissler, OSRD 6487, Report No. 
A-416, Union Carbide & Carbon Research Laboratories, 
Inc., Jan. 17, 1946. Div. 1-420.31-M7 

89. Studies of the Application of Stellite No. 21 to Gun Bores, 

T. H. Gray and D. R. Mosher, OSRD 6488, Report No. 
A-417, Westinghouse Electric and Manufacturing Co., 
Inc., Nov. 12, 1945. Div. 1-420.31-M5 


90. Investigation of Certain Methods for Making Gun Linings 
of Stellite and Other Erosion-Resistant Materials, J. 
Wulff, OSRD 6489, Report No. A-418, MIT, Sept. 28, 

1945. Div. 1-420.31-MI 

91. Preparation and Testing of 37-mm Stellite Liners, J. S. 
Burlew, OSRD 6490, Report No. A-419, CIW. 

Div. 1-420.31-M8 

92. Refractaloy 70 as a Liner Material for Caliber .50 Barrels, 
T. H. Gray, OSRD 6491, Report No. A-420, Westing¬ 
house Electric and Manufacturing Co. Inc., Jan. 10, 

1946. Div. 1-420.35-MI 

93. Pyrolytic Plating from the Carbonyls of Molybdenum, 
Tungsten, and Chromium, L. H. Germer and J. J. Lander, 
OSRD 6492, Report No. A-421, BTL, Nov. 30, 1945. 

Div. 1-630-M2 

94. The Semicommercial Preparation of Molybdenum Car¬ 
bonyl, A. L. McCoy, OSRD 6493, Report No. A-422, 
Climax Molybdenum Co., July 20, 1945. 

Div. 1-420.32-M2 

95. Fabrication of Molybdenum for Use as a Gun-Liner 

Material, J. W. Marden, OSRD 6494, Final Report No. 
A-423, Westinghouse Electric and Manufacturing Co., 
Inc., Oct. 31, 1945. Div. 1-420.32-M3 

96. Development of Molybdenum for Gun Liners, P. H. Brace, 
OSRD 6495, Report No. A-424, Westinghouse Electric 
and Manufacturing Co., Inc., Feb. 1, 1946. 

Div. 1-420.32-M4 

97. Experiments on the Melting of Molybdenum, F. Palmer, 

OSRD 6496, Report No. A-425, Climax Molybdenum 
Co., and Westinghouse Electric and Manufacturing Co., 
Inc., February 1946. Div. 1-420.32-M5 

98. Studies of Erosion Products of Gun-Bore Surfaces, E. G. 
Zies and C. A. Marsh, OSRD 6497, Report No. A-426, 
Geophysical Lab., CIW, Mar. 9, 1945. 

Div. 1-410.1-M13 

99. The Penetration of Carbon into Gun-Bore Surfaces, W. D. 
Urry, E. Jensen, and P. J. Hannan, OSRD 6498, Report 
No. A-427, Geophysical Lab., CIW, Nov. 20, 1945. 

Div. 1-410.1-M9 

100. Work on Sabot-Projectiles by the University of New Mex¬ 
ico under Contract OEMsr-668 and Supplements, 1942- 

1944, J- W. Greig, OSRD 6499, Report No. A-428, 
University of New Mexico, October 1946. 

Div. 1-510.1-M5 

101. Study of Erosion by Adiabatically Compressed Gases, 
W. Garten, Jr., and Gordon L. Davis, OSRD 6500, 
Report No. A-429, Geophysical Lab., CIW, Dec. 21, 

1945. Div. 1-410.1-M12 

102. Statistical Study of Muzzle Erosion Data at the Naval 

Proving Ground, G. Cresson, E. Frankel, and G. Kov- 
sky, OSRD 6501, Report No. A-430, Geophysical Lab., 
CIW, June 21, 1945. . Div. 1-430-MI 

103. An Experiment to Determine the Effects of Stress on 

Gun Erosion, R. W. Goranson and W. Garten, Jr., 
OSRD 6502, Report No. A-431, Geophysical Lab., 
CIW, Nov. 26, 1945. Div. 1-410.3-M2 


CONFIDENTIAL 



BIBLIOGRAPHY 


637 


104. Transient Thermal Action on Gun Steel Induced by 

Electron Bombardment, W. Garten Jr. and Gordon L. 
Davis, OSRD 6503, Report No. A-432, Geophysical 
Lab., CIW, Feb. 15, 1946. Div. 1-410.2-M3 

105. A Method for the Determination, of the Melting Tempera¬ 

tures of Gun Erosion Products, E. Jensen, OSRD 6504, 
Report No. A-433, Geophysical Lab., CIW, Jan. 11, 
1946. Div. 1-410.2-M2 

106. Temperature Distribution in Gun Barrels, G. Comenetz 
and V. Schwab, OSRD 6505, Report No. A-434, 
Geophysical Lab., CIW, March 1946. Div. 1-310-M6 

107. The Quenching of Powder Gas Reactions, F. C. Kracek 
and P. J. Hannan, OSRD 6506, Report No. A-435, Geo¬ 
physical Lab., CIW, Mar. 9, 1946. Div. 1-210.2-M8 

108. Microwave Techniques for Interior Ballistic Measure¬ 
ments, H. S. Roberts, OSRD 6507, Report No. A-436, 
Geophysical Lab., CIW, December 1945. 

Div. 1-210.31-M2 

109. Investigation of Gun Erosion at the Geophysical Laborer 

tory, Volume IV—July to December 1944, P- R. Heyl, 
OSRD 6508, Report No. A-437, Geophysical Lab., 
CIW, February 1946. Div. 1-400-MI 

110. Investigation of Gun Erosion at the Geophysical Labora¬ 

tory, Volume V—January to September 1945, P. R. Heyl, 
OSRD 6509, Report No. A-438, Geophysical Lab., 
CIW, April 1946. Div. 1-400-MI 

111. The Second Approximation for a Yawing Supersonic 
Cone, A. H. Stone, OSRD 6510, Report No. A-439, 
Geophysical Lab., CIW, Nov. 26, 1945. 

Div. 1-220.1-M4 

112. Illustrations and Descriptions of Surface Features of 

Eroded Gun Bores, H. E. Merwin and M. Sullivan, 
OSRD 6511, Report No. A-440, Geophysical Lab., 
CIW, January 1946. Div. 1-400-M4 

113. Comparisons of Interior Ballistic Theory and Experi¬ 
ment — 3-In. and 37-MM Guns Fired at David W. Taylor 
Model Basin, W. S. Benedict, OSRD 6512, Report No. 
A-441, Geophysical Lab., CIW, March 1946. 

Div. 1-210.3-M6 

114. Static Band Pressure in 37-mm Projectiles [Part II], 

F. A. Biberstein, Jr., R. Brown, V. Griffing, K. F. Herz- 
feld, J. M. Krafft, and W. T. Whelan, OSRD 6513, 
Report No. A-442, Catholic University of America, 
Aug. 28, 1945. Div. 1-530-M3 

115. Static Band Pressures in 37-mm Projectiles [Part III], 

F. A. Biberstein, Jr., R. Brown, V. Griffing, K. F. 
Herzfeld, J. M. Krafft, and W. T. Whelan, OSRD 6514, 
Report No. A-443, Catholic University of America, 
December 1945. Div. 1-530-M3 

116. Interior Ballistics of the 3-In. Gun Fired at David W. 
Taylor Model Basin with a Critical Examination of Cer¬ 
tain Assumptions of the Hirschfelder System of Interior 
Ballistics, R. E. Johnson, J. W. Wrench, Jr., O. Kracek, 
and K. F. Herzfeld, OSRD 6515, Report No. A-444, 
Catholic University of America, Nov. 10, 1945. 

Div. 1-210.1-M10 


117. Stresses in Shells Due to Band Pressure, K. F. Herzfeld, 
OSRD 6516, Final Report No. A-445, Catholic Uni¬ 
versity of America, Dec. 28, 1945. Div. 1-530-M4 

118. Design and Construction of Tubes for a Hyper-Velocity 
90-mm Gun, J. H. Billings, OSRD 6539, Report No. 
A-446, Drexel Institute of Technology, February 1946. 

Div. 1-340-M3 

119. Pilot Plant for Production of Modified Caliber .50 Ma¬ 
chine Gun Barrels with Stellite Liners, R. A. Mueller, 
F. D. Cotterman, and J. P. Magos, OSRD 6518, Report 
No. A-447, Crane Co., Oct. 17, 1945. Div. 1-420.31-M4 

120. Pre-Engraved Projectiles, N. H. Smith, OSRD 6519, Re¬ 
port No. A-448, The Franklin Institute, Dec. 6, 1945. 

Div. 1-520-M2 

121. Fisa Protectors — Design, Production, and Tests, N. H. 

Smith, OSRD 6520, Report No. A-449, The Franklin 
Institute, Nov. 20, 1945. Div. 1-420.5-MI 

122. The Caliber .50 Erosion Testing Gun, N. H. Smith, 

OSRD 6521, Report No. A-450, The Franklin Institute, 
Jan. 7, 1946. Div. 1-420.22-MI 

123. Comparison of the Erosiveness of Propellant Powders, 
N. H. Smith, OSRD 6522, Report No. A-451, The 
Franklin Institute, Oct. 12, 1945. Div. 1-420.1-MI 

124. Application of Heat Transfer Theory to Metallographic 
Evidences of Gun Erosion, J. N. Hobstetter, OSRD 6523, 
Report No. A-452, Harvard University, Dec. 20, 1945. 

Div. 1-410.2-MI 

125. Production of Stellite Liners by Centrifugal Casting, W. 
H. Shallenberger, OSRD 6524, Report No. A-453, In¬ 
dustrial Research Laboratories, Inc., December 1945. 

Div. 1-420.31-M6 

126. Development of 20-mm Automatic Aircraft Cannon, W. 

H. Shallenberger, OSRD 6525, Report No. A-454, John¬ 
son Automatics, Inc., and University of California, 
Dec. 18, 1945. Div. 1-610-MI 

127. Production of Modified Caliber .30 Machine-Gun Barrels 

with Stellite Liners, M. M. Johnson, Jr., OSRD 6526, 
Final Report No. A-455, Johnson Automatics, Inc., 
Sept. 28, 1945. Div. 1-420.31-M2 

128. 57/40-mm Tapered-Bore Gun Tubes and Deformable 

Projectiles, OSRD 6527, Report No. A-456, Jones & 
Lamson Machine Co., and Bryant Chucking Co., Oct. 
24, 1945. Div. 1-330-M2 

129. Contributions to the Development of Erosion-Resistant 
Materials for Gun Liners and Linings, P. H. Brace, 
OSRD 6528, Report No. A-457, Westinghouse Electric 
and Manufacturing Co., Inc., June 24, 1946. 

Div. 1-420.3-M3 

130. Studies of Worn Muzzle Sections of Guns by Laboratory 
Techniques, L. E. Line, Jr., J. N. Hobstetter, and Mem¬ 
bers of the Staff of the Geophysical Laboratory, CIW, 
OSRD 6529, Report No. A-458, Harvard University 
and Geophysical Lab., CIW, Jan. 28, 1946. 

Div. 1-430-M3 


CONFIDENTIAL 



638 


BIBLIOGRAPHY 


131. Report on Firings of 37-mm, Gun TJfY, at David W. 
Taylor Model Basin, OSRD 6530, Report No. A-459, 
NBS and Geophysical Lab., CIW, Oct. 22, 1945. 

Div. 1-210.3-M4 

132. Third Report on Ballistic Firings with 3-In. Gun at 
David W. Taylor Model Basin, OSRD 6531, Report No. 
A-460, NBS and Geophysical Lab., CIW, Mar. 25,1946. 

Div. 1-210.3-M5 

133. Development of an All-Metal Type of Sabot-Projectile, 

OSRD 6532, Report No. A-461, Remington Arms Co., 
Oct. 12, 1945. Div. 1-510.1-M3 

134. A Production Process for the Manufacture of an All- 
Metal Type of Sabot-Projectile, OSRD 6533, Report No. 
A-462, Remington Arms Co., Oct. 12, 1945. 

Div. 1-510.1-M4 

135. Production of Modified Machine-Gun Barrels with Stel¬ 
lite Liners Including Studies of Draw Rifling, OSRD 
6534, Report No. A-463, Remington Arms Co., Oct. 3, 

1945. Div. 1-420.31-M3 

136. Hastelloy C as a Liner Material for Machine-Gun Barrels, 

F. S. Badger and W. A. Wissler, OSRD 6535, Report 
No. A-464, Union Carbide and Carbon Research Lab¬ 
oratories, Inc., May 16, 1946. Div. 1-420.34-MI 

137 Electron and X-Ray Diffraction Studies of Gun Erosion 
Products, F. E. Haworth, OSRD 6536, Report No. A- 
465, BTL, Dec. 6, 1945. Div. 1-410.1-M11 

138. Chemical Thermodynamics of Gun Erosion, J. J. Lander, 
OSRD 6537, Report No. A-466, BTL, Apr. 14. 1943. 

Div. 1-410.1-M3 

139. Investigation of the Control of Erosion in Guns and the 
Improvement of Gun Performance, P. R. Heyl, OSRD 
6538, Report No. A-467, Geophysical Lab., CIW, June 

1946, Revised July 31, 1946. Div. 1-400-M5 

140. Ballistic Research, Project Summaries for Division 1 , as 
of October 1, 1945, Report No. A-383, Oct. 8, 1945. 

Div. 1-100-MI 

141. Determination of the Linear Burning Rates of Propellants 
from Pressure Measurements in the Closed Chamber, 
L. G. Bonner, OSRD 4382, Duke University (for Di¬ 
visions 1 and 3, NDRC), Nov. 30, 1944. 

142. Pilot Plant for Hard Chromium Plating Caliber .50 Ma¬ 
chine Gun Barrels, Arnold Weisselberg, Chrome Gage 
Corporation, OSRD 6517, Nov. 16, 1945. 

OTHER REPORTS ISSUED BY OSRD b 

143. Report of the Ad Hoc Committee on Internal Ballistics of 
the National Defense Research Committee, Sept. 29, 1942. 

144. Cranz’s Textbook of Ballistics: Vol. II, Interior Ballistics 
and Vol. Ill, Experimental Ballistics, translated by C. 
C. Bramble and Henry ^luestone, translation revised 
and edited by J. D. Elder and Duane Roller, Technical 


b Items 539 and 547 in Additional References of this Bibliog¬ 
raphy pertain to this section. 


Reports Section, NDRC, May 6, 1944 and Apr. 9, 1945. 
(See Item 505.) Div. 3-220-M3 

145. The Effect of Hyper-Velocity on the Probability of Hitting 
Tanks, E. J. Moulton, W. Weaver, OEMsr-1007 Report 
No. 26.1, Applied Mathematics Panel, July 1, 1943. 

Div. AMP-901.1-MI 

146. Comment on the Memorandum “ Aircraft Machine Guns, 
the Evaluation of Hyper-Velocity Types of, for Naval Air¬ 
craft, Dated March 18, 1943, and Prepared by the Special 
Board of Naval Ordnance,” Memorandum No. 53.1, Ap¬ 
plied Mathematics Panel, NDRC, July 12, 1943. (See 
Item 325 for Navy Memorandum.) 

Div. AMP-504.1-M10 

147. The Shell Design Problem, AMP Report No. 75.1, 
OEMsr-1007, Applied Mathematics Panel, February 

1944. Div. AMP-903.1-M3 

148. A Report of the Kinematic and Dynamic Analysis of the 
Johnson 20-mm Automatic Cannon, Model I, L. M. K. 
Boelter, R. C. Martinelli, F. E. Romie, F. M. Hamaker, 
and N. W. Snyder, University of California (for OSRD 
Engineering Office), January 1945. 

149. Stress Analysis of Breech Mechanism of the Johnson 
20-mm Automatic Cannon, Model III, L. M. K. Boelter, 
F. M. Hamaker, and F. E. Romie, University of Cal¬ 
ifornia (for OSRD Engineering Office), Jan. 1, 1945. 

150. Firing Rate Analysis of the Johnson 20-mm Automatic 
Cannon, Model III, L. M. K. Boelter, F. M. Hamaker, 
N. W. Snyder, and F. E. Romie, University of Califor¬ 
nia (for OSRD Engineering Office), Jan. 1, 1945. 

151. Analysis of Thermal and Pressure Stresses in Caliber .50 
Light Aircraft Machine Gun Barrel, L. M. K. Boelter, 
N. W. Snyder, R. Bromberg, and L. M. Grossman, 
(Report to OSRD), University of California, January 

1945. 

152. Terminal Ballistic Performance of Non-defarming Pro¬ 
jectiles with Special Reference to High Velocities, C. W. 
Curtis, Princeton Technical Memorandum No. 11, Di¬ 
vision 2, NDRC, Mar. 18, 1944. 

153. High Velocity Performance of Tungsten Carbide Pro¬ 
jectiles, Princeton Technical Memorandum No. 17, 
Division 2, NDRC, Mar. 18, 1944. 

154. Plating Trials of 57/ 40 -mm Deformable Projectile, C. W. 
Curtis and R. J. Emrich, Princeton Technical Memo¬ 
randum No. 81, Division 2, NDRC, November 1944. 

155. Transverse Rupture Strength Measurement of Tungsten 
Carbide Cores for 57/40 mm Deformable Projectile, R. J. 
Emrich, Princeton Technical Memorandum No. 95, 
Division 2, NDRC, Jan. 9, 1945. 

156. Terminal Ballistics of Tungsten Carbide Projectiles: Sur¬ 
vey and Nose-Shape Tests, C. W. Curtis, R. J. Emrich, 
and J. R. Sproule, OSRD 4720b, Report No. OTB-7b, 
Feb. 15, 1945. 

157. Terminal Ballistics of Tungsten Carbide Projectiles: Ef¬ 
fect of Carrier, Part I, E. R. Jones, C. W. Curtis, and 
R. J. Emrich, OSRD 5350a, Report No. OTB-12a, 
July 15, 1945. 


CONFIDENTIAL 




BIBLIOGRAPHY 


639 


158. Perforation Limits for Nonshattering Projectile Against 
Thick Homogeneous Armor at Normal Incidence, C. W. 
Curtis and R. L. Kramer, OSRD 6464, Report No. A- 
393, December 1945. 

159. Effect of Pressure and Temperature on the Rate of Burning 
of Double-Base Powders of Different Compositions, W. H. 
Avery and R. E. Hunt, OSRD-1993, NDRC Report No. 
A-225, Jet Propulsion Research Laboratory, Indian 
Head, Md., October 1943. 

160. The Mechanism of Powder Burning, Farrington Daniels 
and collaborators, OSRD-3206, NDRC Report A-243, 
University of Wisconsin, January 1944. 

161. Observations on the Burning of Double-Base Powders, 
B. L. Crawford, Jr., C. Huggett, and J. J. McBrady, 
OSRD-3544, NDRC Report A-268, University of Min¬ 
nesota, April 1944. 

162. Minutes of the Third Meeting of the Sub-Group on Burn¬ 
ing Rates and Related Problems Held at the Ballistic Re¬ 
search Laboratory, Aberdeen Proving Ground, Aberdeen, 
Maryland, April 28-29, 1944, OSRD 3711, Rocket Pro¬ 
pellant Panel of the Joint Committee on New Weapons 
and Equipment. 

163. Rocket Fundamentals, OSRD-3992, Report No. ABL- 
SR4, Section H, Division 3, NDRC, George Washing¬ 
ton University, Dec. 26, 1944. 

164. The Way Muzzle Velocity Affects the Probability of Hit¬ 
ting Aircraft by A-A Fire, Section D-2, NDRC (Memo¬ 
randum), Apr. 1, 1942. 

165. Measurements of Fluid Friction Loss in 0.50 Caliber 
Rifled and Unrifled Gun Barrels, Robert T. Knapp, 
OSRD 2001, California Institute of Technology, July 
16, 1943. 

166. Firing Strains in a 87-MM Field Gun, A. V. deForest, 
Contract OEMsr-155 (Extra Progress Report), Massa¬ 
chusetts Institute of Technology, Feb. 16, 1942. 

167. Heat-Resisting Metals for Gas Turbine Parts ( N-102 ): 

Chromium-Base Alloys, R. M. Parke, OSRD 5044, 
Progress Report M-510, Climax Molybdenum Co., 
May 7, 1945. Div. 18-502.11-MI 

168. Heat-Resisting Metals for Gas Turbine Parts {N-102): 

Chromium-Base Alloys, R. M. Parke and A. J. Herzig, 
OSRD 6547, Final Report M-656, Climax Molybdenum 
Co., Jan. 21, 1946. Div. 18-502.11-M2 

169. Final Report on Heat-Resisting Metals for Gas Turbine 

Parts {N-102), H. C. Cross and W. F. Simmons, OSRD 
6563, Final Report M-636, Battelle Memorial Institute, 
American Brake Shoe and Foundry Co., and others, 
Jan. 21, 1946. Div. 18-502.1-M11 

170. Prevention of Cracking in Gun Tubes, R. F. Mehl, Cyril 
Wells, Irwin Broverman, J. W. Spretnak, C. F. Sawyer, 
and C. C. Busby, OSRD 5383, Final Report M-555, 
Carnegie Institute of Technology, July 30, 1945. 

Div. 18-302.5-M3 

171. Industrial Applications of Chromium Plating: A Review, 
M. Kolodney, OSRD 1074, Advisory Report M-26, 
National Academy of Sciences, Nov. 27, 1942. 

Div. 18-900-MI 


REPORTS ISSUED BY ABERDEEN 
PROVING GROUND 0 

172. Form Factors of Projectiles, H. P. Hitchcock, BRL- 
APG, June 24, 1942, (Revised later by addition of new 
tables.) 

173. The Equation of State of the Powder Gas, J. P. Vinti, Re¬ 
port No. 288, BRL-APG, July 8, 1942 

174. The Effect of Cross Wind on the Yaw of Projectiles, R. H. 
Kent and H. P. Hitchcock, Report A-IV-31, BRL- 
APG, Feb. 24, 1928. 

175. The Rate of Erosion of the Browning Caliber .30 Tank 
Type Machine Gun as Dependent on the Temperature of 
the Gun and Other Factors, N. A. Tolch, Report No. 13, 
BRL-APG, July 8, 1935. 

176. Resistance Functions of Various Types of Projectiles, 
H. P. Hitchcock, Report No. 27, BRL-APG, June 
1935. 

177. Stability Factors of Projectiles, H. P. Hitchcock, Report 
No. 30, BRL-APG (revised) Sept. 30, 1940. 

178. Heating Effects and Endurance Properties of the 105-mm 
A.A. Gun, Ml, as Determined by Rapid Fire Tests, N. A. 
Tolch, Report No. 35, BRL-APG, Feb. 12, 1936. 

179. Heating Effects and Endurance Properties of the 3-In. 
A. A. Gun, M3, as Determined by Rapid Fire Tests, N. A. 
Tolch, Report No. 39, BRL-APG, Mar. 2, 1936. 

180. Roggla’s Equation and its Application to Interior Ballistic 
Problems, R. H. Kent, Report No. 48, BRL-APG, 
(revised) July 3, 1941. 

181. Study of the Heating and Cooling of the Bore of the 75-mm 
Gun, Ml897 E3, W. D. Dickinson, Report No. 49, 
BRL-APG, Apr. 22, 1936. 

182. Heating Effects and Distribution of the Energy of the 
Charge of the Caliber .50, M2 Machine Gun, Heavy 
Barrel, N. A. Tolch, Report No. 61, BRL-APG, 
Oct. 1, 1936. 

183. Heating Effects and Distribution of the Energy of the 
Charge of the Caliber .50 Machine Gun, M2, Water- 
Cooled Type, N. A. Tolch, Report No. 74, BRL-APG, 
Apr. 5, 1937. 

184. An Elementary Treatment of the Motion of a Spinning 
Projectile about its Center of Gravity, R. H. Kent, 
Report No. 85, BRL-APG, Aug. 16, 1937. (See Report 
No. 459 for revision.) 

185. Development of Chrome Plating of Guns, T. K. Vincent, 
Report No. 87, BRL-APG, Nov. 3, 1937. 

186. Heating of Guns, J. R. Lane, Report No. 104, BRL- 
APG, May 20, 1938. 

187. The Cooling of Guns, J. R. Lane, Report No. 146, 
BRL-APG, Apr. 26, 1939. 

188. A Study of Form Factors of Spinning Projectiles, H. P. 
Hitchcock, Report No. 166, BRL-APG, Dec. 8, 1939. 


c The following items in Additional References of this Bibliog¬ 
raphy pertain to this section: 540, 543, 548, 549, 550, 551, 552, 
554, 555, 558, 560, 561, and 562. 


CONFIDENTIAL 





640 


BIBLIOGRAPHY 


189. Drag on Conical Heads, H. P. Hitchcock, Report No. 
240, BRL-APG, July 14, 1941. 

190. The Performance of Sub-caliber Projectiles Compared 
with that of Conventional Types, R. H. Kent, Report No. 
265, BRL-APG, Jan. 9, 1942. 

191. Cooling Corrections for Closed Chamber Firings , R. H. 
Kent and J. P. Vinti, Report No. 281, BRL-APG, Sept. 
7, 1942. 

192. The Motion of the Axis of a Spinning Shell Inside the 
Bare of the Gun, F. V. Reno, Report No. 320, BRL-APG, 
Feb. 27, 1943. (See Item 201 for a treatment of this 
subject based on a different assumption.) 

193. The Effect of Yaw Upon Aircraft Gunfire Trajectories, T. 
C. Sterne, Report No. 345, BRL-APG, June 11, 1943. 

194. Report on the Temperature Dependence of Rocket Be¬ 
havior, J. H. Frazer, J. H. Wiegand, J. E. Mayer, J. W. 
Perry, C. Fennimore, and H. D. Burnham, Report No. 
353, BRL-APG, May 1, 1943. 

195. Firing Strain Measurements on 76-mm Gun, M1E2, V. 
H. McNeilly, Report No. 362, BRL-APG, June 1, 1943. 

196. Simplified Equations of Interior Ballistics, J. E. Mayer, 
Report No. 388, BRL-APG, Aug. 4, 1943. 

197. On the Motion of a Projectile with Small or Slowly Chang¬ 
ing Yaw, J. L. Kelley and E. J. McShane, Report No. 
446, BRL-APG, Jan. 29, 1944. 

198. Evidence of a Stepwise Mechanism of Powder Burning 
from Closed Chamber Firings, H. D. Burnham, Report 
No. 456, BRL-APG, Mar. 28, 1944. 

199. An Elementary Treatment of the Motion of a Spinning 
Projectile about its Center of Gravity, R. H. Kent and 
E. J. McShane, Report No. 459, BRL-APG, Apr. 11, 
1944. (A revision of Report No. 85.) 

200. Erosion in Vent Plugs, J. H. Wiegand, Report No. 520, 
BRL-APG, Jan. 25,1945. (See Report No. 578 for con¬ 
tinuation.) 

201. The Motion of the Axis of a Spinning Shell Inside the 
Bore of the Gun, L. H. Thomas, Report No. 544, BRL- 
APG, May 8, 1945. (See Item 192 for a treatment of 
this subject based on a different assumption.) 

202. A Method for Computation of the Pressure on the Head of 
a Pointed High-Speed Projectile, I. E. Segal, Report No. 
571, BRL-APG, Aug. 31, 1945. 

203. Erosion in Vent Plugs, II. The Effects of Vent Shape and 
of Metal, J. H. Wiegand, Report No. 578, BRL-APG, 
Jan. 25, 1946. 

204. The Dependence of Muzzle Velocity on the Potential and 
the Force of the Powder, J. P. Vinti and J. Chernick, 
Report No. 583, BRL-APG, Oct. 26, 1945. 

205. Notes on Increasing the Penetration of the Medium Tank 
Gun, S. J. Zaroodny, Memorandum Report No. Ill, 
BRL-APG, Dec. 26, 1941. 

206. Yaw of 3-In./57-mm Sabot Type Projectile Fired from 
British 17-Pounder Gun, H. P. Hitchcock, Memoran¬ 
dum Report No. 171, BRL-APG, June 1, 1943. 


207. Ballistics of NDRC 20-mm Sabot Type Projectiles, A. P. 
Alexander and H. P. Hitchcock, Memorandum Report 
No. 251, BRL-APG, Nov. 27, 1943. 

208. The Observed and Computed Deflections of Bullets Fired 
Sidewise from an Airplane, T. E. Sterne, Memorandum 
Report No. 273, BRL-APG, Feb. 19, 1944. 

209. Designs of Sub-Caliber Projectiles, H. P. Hitchcock, 
Memorandum Report No. 302, BRL-APG, June 17, 
1944. 

210. Flash Radiographs of 76-mm Sabot-Projectile, J. C. Clark, 
Memorandum Report No. 330, BRL-APG, Sept. 14, 

1944. 

211. Effects upon the Moment and Drag Coefficient of an In¬ 
crease in Width of the Driving Band, W. A. Siljander, 
Memorandum Report No. 365, BRL-APG, Apr. 30, 

1945. 

212. Experimental Subcaliber A. P. Projectile for the 37-mm 
Aircraft Gun, M/, S. J. Zaroodny, Report No. 323, 
BRL-APG, Jan. 8, 1943. 

213. Test of Aircraft Machine Gun Barrels, Caliber .50, with 
Stellite Liners, 1st and 3rd Memorandum Reports on 
F. R. No. S-42589, Ordnance Program No. 5082, APG, 
Sept. 14, and Nov. 11, 1944. 

214. Test of Chromium Plated Caliber .50 Aircraft Machine 
Gun Barrels, Lot No. 2, 4th Memorandum Report on 
Firing Record S-42762, Ordnance Program No. 5082, 
APG, Mar. 16, 1945. 

215. Test of Gun, Machine, Caliber .60, T17E3, Aircraft 
[Using Chromium-Plated Barrels Furnished by Division 
One], 12th, 13th, and 14th Memorandum Reports on 
Firing Record S-42900, Ordnance Program No. 5082, 
APG, Apr. 24, 25, 26, 1945. 

216. First Report in Connection with the Test of the Accuracy 
Life of the 155-mm Gun, M1A1E1, No. 3052; and Thirty- 

first Report on 155-mm Gun — 8-inch Howitzer Materiel, 
Ordnance Program No. 5084, (2 Vols), APG, Sept. 6, 
1944. 

217. Test of Special High-Velocity “Hour Glass” Type 37/28- 
mm Projectile by NDRC, Memorandum Report On 
Ordnance Program No. 5364, APG, July 11, 1942. 

218. The Second Report on the 37/28-mm and 37-mm High 
Velocity Armor Piercing Projectiles and Twenty-second 
Report on Ordnance Program No. 536/., APG, July 19, 
1942. 

219. Form Factor, Yaw Recovery, and Breech Pressure Firings 
of 37/28-mm NDRC ( N.B.S. Type G) Projectiles, Firing 
Record No. S-20960, Ordnance Program No. 5364, 
APG, Feb. 3, 1943. 

220. First Report in Connection with the Test of the Accuracy 
Life of Tubes, 90-mm Gun, M3E2 ( Chromium-Plated ) 
and Sixth Report on Ordnance Program No. 5/12, APG, 
Apr. 27,1945. 

221. Test of 37-mm Tube, Stellite Liners A, B, and C, 3rd 
Memorandum Report, Firing Records Nos. M-46044, 
M-46045, M-46269, Ordnance Program No. 5426, APG, 
Sept. 9,1945. 


CONFIDENTIAL 



BIBLIOGRAPHY 


641 


222. First Report on Test of Projectile, Hyper-Velocity Armor 
Piercing (Sabot), 90-mm T32 and T38 Series, First Re¬ 
port on Ordnance Program No. 5757, and First Report on 
Ordnance Program No. 5962, APG, Oct. 2, 1945. 

223. First Report on the Projectile H.V.A.P. (Sabot), 76-mm 
3-76EH (University of New Mexico), T23 and T23EI 
(Remington Arms Company), and Second Report on 
Ordnance Program No. 5962, APG, Oct. 2, 1945. 

224. Firing Test to Determine Ballistic and Functioning Char¬ 
acteristics of 105-mm/3-inch and 105-mm/57-mm Sabot- 
Projectiles (University of New Mexico Design) Firing 
Record No. P-31518, Ordnance Program No. 5762, 
APG, June 7, 8, 1944. 

225. Second and Final Report on 8.8 cm Anti-Tank Gun Pak 
43/41, German, Memorandum Report on Ordnance 
Program No. 5772, APG, Aug. 3, 1945. 

226. Data Relative to the Erosion Resisting Quality of Chro¬ 
mium-plated Caliber .50 Aircraft Barrels, Firing Record 
No. S-33346, Ordnance Program No. 5800, APG, Dec. 
22, 1943 to Feb. 4, 1944. 

227. Test of 37/28-mm Hour Glass Projectiles by NDRC, Fir¬ 
ing Record No. S-14020, Ordnance Program No. 5829, 
APG, July 6 to Nov. 14, 1942. 

228. Practicability of Design of NDRC 37/28-mm Projectiles, 
Firing Record No. S-18216, Ordnance Program No. 
5829, APG, Dec. 15, 1942. 

229. Test of 37/28-mm NDRC Projectiles (N.B.S. Type G 
with Various Deviations), Memorandum Report on 
Ordnance Program No. 5829, APG, Mar. 22, 1943. 

230. Firing Test to Determine by Spark and Micro-Flash 
Photography which of Two 37/28-mm NDRC (NBS Type 
G with .238" and .100" Base Skirts) Projectiles has the 
Better Contour Characteristics in Flight, Firing Record 
No. S-23635, Ordnance Program No. 5829, APG, Apr. 

14.1943. 

231. Tests of Replaceable Liner for 75-mm Gun, M3, First 
through Thirteenth Memorandum Reports on Ord¬ 
nance Program No. 5834, APG, Apr. 22, 1943 to May 

25.1944. 

232. Firing Test to Determine Stability of 3" Sabot-Type Pro¬ 
jectile when Fired from 3" Gun, M1918M1, at Chamber 
Pressure oj 40,000 p.s.i., Firing Record No. P-20826, 
Ordnance Program No. 5867, APG, Apr. 14, 1943. 

233. Firing Test to Determine Suitability and Stability of Ex¬ 
perimental 3" Sabot-Type Projectiles, Firing Record No. 
P-21895, Ordnance Program No. 5867, APG, Apr. 29, 

30.1943. 

234. Determination of Pressure and Velocity Relationship of 
20 mm Subcaliber Projectiles for NDRC, Firing Record 
No. S-22190, Ordnance Program No. 5945, APG, Mar. 

10.1943. 

235. Firing Tests of 3-in./57mm Sabot-Projectile (Critchfield 
Design) Fired in British 17-Pounder (3-in.) Gun, Mk. I, 
Firing Records Nos. M-23473 (Ordnance Program No. 
5954) and M-22798 (Ordnance Program No. 5965), 
APG, May and June 1943. 


236. Firing Tests of 105/75-mm Sabot-Projectile, Critchfield 
Design, Firing Records Nos. M-23342 and M-24905. 
Ordnance Program No. 5954, APG, June and July 1943. 

237. Stability Firing Program on the 20-mm Sabot-Type Pro¬ 
jectile for the Ballistic Research Laboratory, Firing Rec¬ 
ord No. S-31707, Ordnance Program No. 5962, APG, 
June 7, Sept. 8, 9, 25, 27, and Oct. 5, 1943. 

238. Firing Tests of 75/57-mm Sabot-Type Projectiles, Uni¬ 
versity of New Mexico, Design 28-75D, Firing Records 
Nos. M-24862, M-25496, M-28120, and M-28827, Ord¬ 
nance Program No. 5962, APG, July 12 to Dec. 23, 
1943. 

239. Firing Test to Obtain Micro-Flash Pictures of 75-mm 
Eksergian Sabot-Type Projectiles in Flight, Firing Rec¬ 
ord No. M-24957, Ordnance Program No. 5962, APG, 
July, 19, 21, 1943. 

240. Proof Firing of 37-mm Tubes, T47 and Establishment of 
Charge for 37-mm Experimental Pre-Engraved Projectiles 
in 37-mm Gun, T47, Firing Record No. P-37989, 
Ordnance Program No. 5996, APG, June 6 to 21, 1945. 

241. First Report on Development Test of 90-mm Gun, T19, 
with a Removable Liner and First Report on Ordnance 
Program No. 6102, APG, July 9, 1945. 

242. Test of 3-in./57-mm Sabot-Type Projectile (Critchfield 
Design) Fired Against Armor, Projectile Test, Report 
No. AD-P99, APG, Aug. 29 to Nov. 18, 1943. 

243. The Action of p-Phenylenediamine on Nitrocellulose, C. 
P. Fennimore, Report No. 612, BRL-APG, June 25, 
1946. 

REPORTS ISSUED BY WATERTOWN ARSENAL d 

244. Report of Conference on Gun Erosion, WA, Watertown, 
Mass., Oct. 15, 1941. 

245. Erosion in Gun Barrels (Monthly Progress Reports to 

Watertown Arsenal on War Department Contract), 
H. W. Russell, Battelle Memorial Institute, February 
12, 1941 to June 30, 1942, pages 1 to 430, inch; July 1, 
1942 to December 31, 1942, pages 1 to 82, inch; January 
31, 1943 to-. 

246. Metallurgical Tests of German 8.8 cm. Gun, First, Sec¬ 
ond, and Third Partial Reports, WA File Nos. 386.3/22, 
386.3/23, 386.3/28, WA, May 14, 27, June 12, 1943. 

247. Metallurgical Examination of Three 76-mm Gun Tubes, 
Two After Endurance Firing and One After Proof Firing, 
N. L. Reed, Experimental Report No. WAL 730/56, 
WA, July 17, 1944. 

248. Examination of 76-mm Gun Tube Ml, No. 52, which was 
Fired 2220 Rounds at Aberdeen Proving Ground, A. M. 
White, Experimental Report No. WAL 730/65, WA, 
Nov. 13, 1943. 

249. Stresses in Gun Tubes: Band Pressure Characteristics of 
37-mm, M3 Gun No. 770, Tube No. 2928, R. Beeuwkes, 


d Items 541, 557, and 563 in Additional References of this 
Bibliography pertain to this section. 


CONFIDENTIAL 





BIBLIOGRAPHY 


642 


Jr., Experimental Report No. 730/95, WA, Apr. 21, 

1944. 

250. Strains in Gun Tubes: Calculation of Pressure Expansion 
Curves of Circular Cylinders , R. Beeuwkes, Jr., J. H. 
Laning, Jr., Experimental Report No. WAL 730/111, 

WA, Mar. 8, 1944. 

251. Stresses in Guns under Combined Band and Gas Pres¬ 
sures, J. H. Laning, Jr., Memorandum Report No. 

WAL 730/118, WA, Mar. 25, 1944. 

252. Firing Strains in Guns Arising from Passive Resistances, 

J. H. Laning, Jr., Memorandum Report No. WAL 
730/123, WA, Apr. 5, 1944. 

253. Stresses in Guns under Combined Band and Gas Pres¬ 
sures, Part 1, J. H. Laning, Jr., Experimental Report 
No. WAL 730/137, WA, Apr. 28, 1944. 

254. Stresses in Guns Under Combined Band and Gas Pres¬ 
sures. Part 2. Elastic Strains at the Outer Surf are for In¬ 
ternal Radial Pressures, Basic Data, O. L. Bowie, J. H. 
Laning, Jr., Experimental Report No. WAL 730/137-1, 

WA, Mar. 22, 1945. 

255. Stresses in Guns under Combined Band and Gas Pres¬ 
sures. Part 3. Elastic Strains at the Outer Surface for In¬ 
ternal Frictional Forces, Basic Data, O. L. Bowie, J. H. 
Laning, Jr., Experimental Report No. WAL 730/137-2, 

WA, May 18, 1945. 

256. Stresses in Guns under Combined Band and Gas Pres¬ 
sures. Part 4- Maximum Equivalent Tensile Stress at the 
Bare Surface, O. L. Bowie, J. H. Laning, Jr., Experi¬ 
mental Report No. WAL 730/137-3, WA, Aug. 31, 1945. 

257. Centrifugal Casting: Metallurgical Examination of 76-mm 
M 1 A 1 Gun Tube No. 1025 After Extended Firing Test, 

P. A. G. Carbonare, Experimental Report No. WAL 
730/144, WA, Oct. 5, 1944. (See Appendix, Report No. 

WAL 731/69-5, Item 267.) 

258. Methods far Computation of Plastic Stresses in Guns 
Under Combined Band and Gas Pressures, J. H. Laning, 

Jr., Memorandum Report No. WAL 730/166, WA, 

July 6, 1944. 

259. Metallurgical Examination of 76-mm Gun Tubes M1A1 
No. 1441 and MIE 6 N 0 .1033 after Extended Firing Tests 
and Detonation Tests at Aberdeen Proving Ground, A. M. 
Burghardt, Experimental Report No. WAL 730/197, 

WA, Oct. 24, 1944. 

260. Erosion II. Plug Erosion Tests, P. R. Kosting, Report 
No. 731/37, WA, June 15, 1938. 

261. The Effect of Service upon 75-mm Gun M1897 No. 1537, 

P. R. Kosting, Report No. 731/44, WA, Aug. 29, 1939. 

262. 37-mm Browning Automatic Gun Ml, No. 1 Barrel, P. R. 277. 
Kosting, Report No. 731/45, WA, Oct. 17, 1940. 

263. Erosion of Guns: Chromium Plated 37-mm Gun Tube 

Ml, No. 4 ( C-4557), R. E. Peterson, Report No. 731/ 278. 

63, WA, Dec. 21, 1942. 

264. Erosion and Progressive Stress-Damage in 76-mm Gun 

Tube Ml, No. 341, P. R. Kosting, Memorandum Re- 279. 
port No. WAL 731/69, WA, Nov. 15, 1943. 


265. The Extent of Progressive Stress-Damage and of Erosion 
in 76-mm Gun Tube M 1 A 1 , No. 1425, P. R. Kosting 
and R. H. Patterson, Memorandum Report No. WAL 
731/69-3, WA, Jan. 14, 1944. 

266. Erosion and Progressive Stress-Damage in 76-mm Gun 
Tubes, Including T 1 No. 2, P. R. Kosting, Memo¬ 
randum Report No. WAL 731/69-4, WA, July 8, 1944. 

267. Progressive Stress-Damage and Erosion in Centrifugally 
Cast and Cold Worked 76-mm Gun Tube M1A1 No. 1025, 
P. R. Kosting and J. A. Kornatowski, Memorandum 
Report No. WAL 731/69-5, WA, Sept. 11, 1944. (Ap¬ 
pendix to Report No. WAL 730/144, Item 257.) 

268. Spectroscopic Analysis of Bore Interface of 76-mm Gun 
Tubes Ml No. 52 and M 1 A 1 No. 1425, P. R. Kosting, 
Memorandum Report No. 731/69-6, WA, Nov. 7, 1944. 

269. Erosion and Progressive Stress-Damage of Guns: 76-mm, 
M 1 E 2 , Tube No. 1105, P. R. Kosting and R. H. Patterson, 
Memorandum Report No. 731/69-7, WA, Dec. 18, 1944. 

270. Erosion of Metals, Alloys, Nitrided Steel, and Metallic 
Coatings on Steel, P. R. Kosting, Experimental Report 
No. WAL 731/72-1, WA, Mar. 24, 1944. 

271. The Chromium Plating and Firing Tests on Seven 37-mm, 
M3 Gun Tubes, (Report to Watertown Arsenal on War 
Department Contract), H. W. Russell, L. R. Jackson, 
E. J. Ramaley, and C. L. Faust, (Watertown File No. 
731/75-8), Battelle Memorial Institute, Feb. 15, 1944. 

272. Memorandum Concerning Chromium Plating for Erosion 
Resistance by Battelle Memorial Institute, L. R. Jackson, 
Battelle Memorial Institute, Report No. WAL 731/ 
75-12, WA, Apr. 14, 1944. 

273. Comparison of Rotating Bands of 37-mm, 75-mm, 3-inch, 
90-mm, 105-mm, and 155-mm Projectiles, P. R. Kosting 
and H. E. Squires, Memorandum Report No. WAL 
731/90-1, WA, Jan. 26, 1945. 

274. Erosion and Progressive Stress-Damage of Guns: 90-mm, 
Ml, Tube No. 7, P. R. Kosting, Memorandum Report 
No. WAL 731/90-2, WA, Jan. 29, 1945. 

275. Erosion and Progressive Stress-Damage of Cannon, M. 
Flynn, R. H. Patterson, and P. R. Kosting, Memoran¬ 
dum Report No. WAL 731/90-4, WA, Dec. 13, 1945. 

276. The Chromium Plating of Six 37-mm Gun Tubes and 
Four Special 15-inch Sections of 37-mm Gun Tubes, 
(Report to Watertown Arsenal on War Department 
Contract), L. R. Jackson, C. L. Faust, and E. L. Combs, 
(Watertown File No. 731/94-1), Battelle Memorial In¬ 
stitute, Nov. 1, 1944. 

Progressive Stress-Damage of Gun Tubes, First Partial 
Report, P. R. Kosting, Experimental Report No. WAL 
731/95, WA, May 2, 1944. 

Progressive Stress-Damage of Gun Tubes, Second Partial 
Report, P. It. Kosting, Experimental Report No. WAL 
731/95-1, WA, Oct. 2, 1944. 

Film in Chromium Electroplate, J. B. Cohen, Experi - 
mental Report No. WAL 731/99, WA, June 3, 1944. 


CONFIDENTIAL 




BIBLIOGRAPHY 


643 


280. Progressive Stress-Damage in 37-mm, MS Tubes, No. 
1500 and No. J+01+0, J. A. Kornatowski and P. R. Kost- 
ing, Experimental Report No. WAL 731/102, WA, 
May 30, 1944. 

281. Progressive Stress-Damage in 37-mm, M6 Tube No. 
102770 ( NDRC-CIT Test), J. A. Kornatowski and P. R. 
Kosting, Experimental Report No. WAL 731/111, WA, 
July 11, 1944. 

282. The Erosion of 37-mm M1A2 Gun Tubes, Including the 
Effect of Propellant Powders FNH-M1 and FNH-M2 
and of Band Diameter, P. R. Kosting, Experimental 
Report No. 731/115, WA, Aug. 7, 1944. 

283. Investigation of 75-mm Gun Tube T13E1 No. 36209 
After 3608 Rounds at Normal Pressure, P. R. Kosting 
and J. A. Kornatowski, Experimental Report No. WAL 
731/126, WA, Nov. 17, 1944. 

284. Investigation of 155-mm Gun Tube M1A1 No. 98 Which 
was Returned from the European Theater of War, P. R. 
Kosting and Mildred D. Flynn, Experimental Report 
No. WAL 731/133, WA, Dec. 22, 1944. 

285. A Detailed Study of the Bore Interface of Chromium 
Plated Tube No. 1521 from 155-mm Gun M1A1E1 No. 
3052 after 2353 Rounds, J. B. Cohen and J. A. Korna¬ 
towski, Experimental Report No. WAL 731/173, WA, 
Sept. 11, 1944. 

OTHER REPORTS ISSUED BY THE WAR 
DEPARTMENT e 

286. High and Low Temperature Ballistic Research: First 
Progress Report on Firings in 76-mm Gun, Ml, Techni¬ 
cal Division, OCO, July 3, 1943. 

287. Stability and Resistance Firings, Ordnance Proof Man¬ 
ual No. 15-10, OCO, Aug. 7, 1943. 

288. Proof Technique: Inspection of Cannon and Measure¬ 
ment of Significant Dimensions, Ordnance Proof Manual 
No. 40-15, OCO, Feb. 26, 1943. 

289. Velocity Measurements of Projectiles, Ordnance Proof 
Manual No. 40-17, OCO, Dec. 21, 1942. 

290. Steel Forgings for Cannon Tubes, U. S. Army Specifica¬ 
tion No. 57-106A, War Department Manual, Jan. 1, 
1945. 

291. Cobalt-Chromium-Alloy; Castings, Investment, Ordnance 
Department Tentative Specification, AXS-1494, Jan. 
11,1945. 

292. Procedure and Final Detail Physical, Dimensional, and 
Performance Requirements for Nitrided and Chromium 
Plated Caliber .50 AC Machine Gun Barrels, 2nd Ed., 
OCO, Feb. 10, 1945. 

293. Design Procedure for Cannon Tubes, Memorandum No. 
7, Cannon Branch, Armament Development Division, 
Research and Development Service, OCO, 1945. 


e Items 538, 545, 546, and 553 in Additional References of this 
Bibliography pertain to this section. 


294. Barrel, Gun, Machine, Caliber .50, Improvement in 
Accuracy Life—Status of Project for, Ordnance Commit¬ 
tee Minute 26230, OCO, (Approved) Jan. 4, 1945. 

295. Barrel (D-716181 If) for Gun, Machine, Browning, Caliber 
.50, M2, Heavy Barrel—Recommended for Adoption, 
Ordnance Committee Minute 27817, OCO, (Approved) 
May 31, 1945. 

296. Barrels, Machine Gun, with Chromium-plated Bores — 
Change in Security Classification Recommended, Ord¬ 
nance Committee Minute 28357, OCO, (Approved) 
July 12, 1945. 

297. Barrel (D-7162295) for Gun, Machine, Caliber .30, 
Browning, M1919A6—Recommended for Adoption, Ord¬ 
nance Committee Minute 28510, OCO, (Approved) 
July 26, 1945. 

298. Barrel ( D-7162295) for Gun, Machine, Caliber .30, 
Browning, M1919A6—Recommendation for Adoption 
Approved; Barrel (D-716181 It) for Gun, Machine, Brown¬ 
ing, Caliber .50, M2, Heavy Barrel—Recommendation for 
Adoption Approved, Ordnance Committee Minute 28893, 
(read for record) OCO, Aug. 30, 1945. 

299. Test of Stellite Lined Caliber .50 Machine Gun Barrels, 
Report on Project No. 4238C472.91, Army Air Forces 
Board, Orlando, Fla., June 28, 1945. 

300. Bofors 57-mm A-A Gun, Military Attache Stockholm 
Report No. 1421, Military Intelligence Division, War 
Department, Oct. 30, 1941. 

301. Penetration Range Curves for 6 Pr. Projectiles, A.R.D./ 
T.H.V. N0.-218, Inclosure to Military Attache London 
Report No. 64056, Military Intelligence Division, War 
Department, Dec. 29, 1943. 

302. N.A.F.D. U. Cooling System for Cal. .50 Aircraft Machine 
Guns, Military Attache London Report No. R174-45, 
Military Intelligence Division, War Department, Jan. 
8, 1945. 

303. Report on 28-cm Projectile, G-2 Report 196, File No. 0.0. 
386.3/1542, Incl. 1, OCO, Apr. 12, 1944. 

304. German 28-cm Railway Gun, H. M. Loewy, Ordnance 
Intelligence Unit (D), APO 464, U. S. Army, June 12, 
1944. 

305. German Sabot Projectiles, H. L. Karsch and Purcell, 
Ordnance Technical Intelligence, Hq. Com Z, USFET. 

306. Chrome Plating of Barrels by the Heinrich Reining, G.m. 
b.H., J. M. Crews, A. E. Kramer, CIOS Target Nos. 
2/722 and 21/503, Artillery and Weapons Metallurgy, 
Combined Intelligence Objectives Sub-Committee, G-2 
Division, SHAEF (Rear) APO 413, June 1945. 

307. Radial Vibrations of Gun Tube, G. C. Evans, T.D.B.S. 
Report No. 15, OCO, Sept. 22, 1943. 

308. Use of Tables to Calculate Band Pressures from Strain 
Gauge Data, G. C. Evans, T.D.B.S. Report No. 30, 
OCO, Apr. 4, 1944. 

309. Criteria for Dimensions of Rotating Bands, G. C. Evans, 
T.D.B.S. Report No. 32, OCO, June 7, 1944. 


CONFIDENTIAL 




BIBLIOGRAPHY 


644 


310. Notes on Gun Design, G. C. Evans, T.D.B.S. Report 
No. 36, OCO, June 5, 1944. 

311. Engineering Report on Chromium Plating Various Small 
Arms Barrels, A. Willink, Frankford Arsenal, July 28, 
1931. 

312. Caliber .50 Interior Ballistic Data, E. R. Thilo, Report 
No. T-1401, Ordnance Laboratory, Frankford Arsenal, 
Nov. 25, 1942. 

313. Development of a Standard Erosion Gage Weapon and a 
Standardized Erosion Test, W. J. Kroeger and C. W. 
Musser, Report No. R-268, 1st Report, Research Item 
1004.0, Ordnance Laboratory, Frankford Arsenal, 
January 1943. 

314. An Accelerated Erosion Test in the Erosion Gage, W. J. 
Kroeger and S. Fernbach, Report No. R-269, 2nd Partial 
Report, Research Item No. 1004.0, Ordnance Labora¬ 
tory, Frankford Arsenal, January 1943. 

315. A Reproducibility Test of the Caliber .30 Erosion Gage, 
W. J. Kroeger and S. D. Rolle, Report No. R-331, 
3rd Partial Report, Research Item 1004.0, Ordnance 
Laboratory, Frankford Arsenal, May 1943. 

316. Determination of Quickness of Standard Pyro and FNH 
Powders, Technical Reports Nos. 966, 1017, 1050, 
Chemical Department, Technical Group, Picatinny 
Arsenal, May 17, 1939, Jan. 3, 1940, and July 26, 1940 

317. Thermochemical and Physical Tests of Nitro-guani- 
dine Powders, Technical Report No. 1336, Chemical 
Department, Technical Group, Picatinny Arsenal, 
Sept. 6,1943. 

318. Investigation of Erosion in Caliber .50 M2 AC BMG 
Barrels, Partial Technical Reports on S.A. W.O. 311, 
Engineering Department, Experimental Division, 
Springfield Armory, Nov. 12, 1943, Mar. 1, 1944, and 
June 6, 1944. 

319. Reports on Tests of Special Aircraft Caliber .50 Machine- 
Gun Barrels (Prepared by Division One, NDRC ), Re¬ 
ports Nos. 29, 34, 36, 37, 42, 47, 48, 50, 52, 54, 58, 60, 
63, 64, 65, 66, 67, 68, 75, 76, 77, 81,82, 87,94, 95, 98, 108 
(each report has an individual title), Purdue University, 
Engineering Experiment Station, Small Arms Ordnance 
Research, (under contract with War Department), 
August 1944—May 1945. 

320. Reports on Tests of Special Caliber .30 Machine-Gun 
Barrels (Prepared by Division One, NDRC), Reports 
Nos. 72, 77, 91, 102, 103, 107 (each report has an in¬ 
dividual title), Purdue University, Engineering Experi¬ 
ment Station, Small Arms Ordnance Research (under 
contract with War Department), January 1945—May 
1945. 

321. [Apparatus for Measuring Bore-Surface Temperatures], 
Report 26B, Purdue University, Engineering Experi¬ 
ment Station, Small Arms Ordnance Research (Final 
Report Covering the Artillery Research Work under 
Contract W-11-022-ORD-4266), June 30, 1946. 


REPORTS ISSUED BY NAVY DEPARTMENT 

322. Partial Report on Chromizing of Steels, NRL Report No. 
M-1822, Navy Department, Dec. 16, 1941. 

323. Reports from Special Board on Naval Ordnance, October 1 
and 15, 1926, Concerning 3-in./105-Caliber Gun, quoted 
in letter EDl(Re5a) from Capt. G. L. Schuyler to J. S. 
Burlew, NDRC, Mar. 28, 1942. 

324. “ Probert” Gun and Projectile Design, Naval Attache 
London Report Serial No. 838, Naval Intelligence Divi¬ 
sion, Navy Department, April 1942. 

325. Aircraft Machine Guns, the Evaluation of Hyper-Velocity 
Types of, for Naval Aircraft, Memorandum from the 
Special Board on Naval Ordnance, Navy Department, 
Mar. 18, 1943. 

326. Thermal Deformation of Gun Bores, C. W. MacGregor 
and L. F. Coffin, Jr., Report to the Chief of the Bureau 
of Ordnance, Navy Contract No. NOrd-470, MIT, Sept. 
20, 1943. 

327. The Effect of Projectile Bands on the Strength of Gun 
Barrels, C. W. MacGregor and L. F. Coffin, Jr., Report 
to the Chief of the Bureau of Ordnance, Navy Contract 
No. NOrd-470, MIT, Sept. 20, 1943. 

328. Equipment and Procedures for Chrome Plating 3"/50 
Cal., 5"/25 Cal., 5"/3$ Cal., 6"/47 Cal., 5"/51 Cal., 
6"/53 Cal., 8"/55 Cal., 12"/50 Cal., U”/45 Cal., 11+"/50 
Cal., 16"/45 Cal., and 16"/50 Cal. Guns, Informal 
Report, Naval Gun Factory, 1943. 

329. 20-mm Aircraft Automatic Gun (Design), Navord Ord¬ 
nance Specifications, NAVORD OS 3380, Navy De¬ 
partment, June, 23, 1944. 

330. Strain Gauge Measurements on 3"/50 Guns, Report No. 
5-45, U. S. Naval Proving Ground, Dahlgren, Virginia, 
Mar. 20, 1945. 

331. Supersonic Flow around Yawing Cones, Z. Kopal, MIT, 
Report No. 186 (Informal) to Bureau of Ordnance, 
Navy Department, 1945. 

332. Technical Conference on Supersonic Flow and Shock 
Waves, Z. Kopal, NAVORD Report No. 203-45, Navy 
Department, 1945. 

REPORTS ISSUED BY BRITISH MINISTRY 
OF SUPPLY 1 ’ g 

333. Temperature Measurements in Guns, AC 1099/Gn.38, 
Met. Report 38 3/41, Aug. 11, 1941. 

334. The Structure of the Reaction Zone in the Burning of Col¬ 
loidal Propellant, S. F. Boys and J. Corner, AC 1139/ 
IB.8, (WA-66-49), August 1941. 


following the abbreviations AC, ADD, ARD, and RD, 
the file number of the report in the OSRD Liaison Office is 
given. 

g Item 556, in Additional References of this Bibliography 
pertains to this section. 


CONFIDENTIAL 




BIBLIOGRAPHY 


335. Thermal Decomposition of S. C. Cordite (Interim Re¬ 
port), C. E.H. Bawn and R. F. J. Freeman, AC 1398/ 
IB.42, (W-112-8), November 1941. 

336. Note on Application of Results Obtained by Relaxation 
Methods , F. Smithies, AC 1553/Gn.81, (II-5-842), Jan¬ 
uary 1942. 

337. Preliminary Trial to Establish the Possibility of Electric 
Communication with the Projectile along the Bore , A. J. 
Garratt, AC 1578/Gn.86 (RD/Ball. Report 8/41), 
(II-5-1036), January 1942. 

338. Further Note on Application of Results Obtained by Re¬ 
laxation Methods , F. Smithies, AC 1648/Gn.91, (OB-1- 
15-1), January 1942. 

339. Strains in a Gun Barrel near the Driving Barrel [Band] of 
a Moving Projectile , G. I. Taylor, AC 1851/GN.104, 
(II-5-1065), March 1942. 

340. Thermochemical Data for the Products of Propellant Ex¬ 
plosion, AC 1862 IB/FP.20, (WA-252-52), March 1942. 

341. Preliminary Calculations on Stresses in Shell, R. Y. 
Southwell, AC 2201/Gn.l27, (WA-177-43), June 1942. 

342. The Measurement of Shot Resistance in the Bore of a Gun, 
R. E. Kutterer, AC 2244/IB.107, (WA-183-25). 

343. Stresses in H. E. Shell: Results Obtained by Relaxation 
Methods, 2nd Report, AC 2457/Gn.l53, (WA-259-47), 
August 1942. 

344. Stresses in Shell: Comparison of Approximate and Re¬ 
laxation Methods, D. G. Sop with, AC 2680/Gn.l70 
(WA-260-15), September 1942. 

345. The Thermal Sensitivity of Explosives, Seventh Report, 
A. J. B. Robertson, AC 2789/S.E.76, (WA-315-10), 
October 1942. 

346. Application of Relaxation Methods to Calculation of 
Stresses in Shell, R. V. Southwell, AC 2797/Gn.l80, 
(WA-315-3), October 1942. 

347. Stresses in H. E. Shell: Comparison of Approximate and 
Relaxation Methods —II, D. G. Sopwith, AC 3066/Gn. 
198, (WA-384-12), November 1942. 

348. The Heating of a Gun Barrel by the Propellant Gases, 
E. P. Hicks and C. K. Thornhill, AC 3119/IB.146, 
(WA-507-1), December 1942. 

349. Stresses in Shell, AC 3136/Gn.200, (WA-384-18), De¬ 
cember 1942. 

350. A Method for the Oscillographic Study of Rapid Pressure 
Changes; Thermal Decomposition of R.D.X. and Ignition 
of Nitrocellulose, A. J. B. Robertson, Eighth Report on 
the Thermal Sensitivity of Explosives, AC 3264/SE.108, 
(WA-436-11), Dec. 22, 1942. 

351. Stresses in Shell, D. N. de G. Allen, AC 3276/Gn.208, 
(WA-403-5), December 1942. 

352. Spectrum of Cordite Burning in an Inert Atmosphere, 
(Interim Report), W. C. Price and R. W. G. Norrish, 
AC 3410/IB.FP.133, (WA-334-14), January 1943. 

353. The Methods of Internal Ballistic Calculation in Use in 
the Armament Research Department, Part 1. The Present 


645 


Working Method, S. W. Coppack, AC 3605/IB. 161 
(ARD/Ball. Report 8^/42, December 1942), (WA- 
532-6), March 1943. 

354. The Method? of Internal Ballistic Calculation in Use in 
the Armament Research Department. Part 2. Some De¬ 
velopments towards Greater Theoretical Completeness, 
S. W. Coppack, AC 3606/IB.162, (ARD/Ball. Report 
82/42, December 1942), (WA-532-7), March 1943. 

355. Stresses in Shell, D. N. de G. Allen, AC 3632/Gn.236, 
(WA-542-13), March 1943. 

356. Stresses in Shells, AC 3867/Gn.246, (WA-620-11), April 
1943. 

357. Stresses in Shell, D. N. de G. Allen, AC 3883/Gn.247, 
(WA-620-12), April 1943. 

358. A Spectroscopic Study of the Light Emitted by Cordite 
Burning in Closed and Vented Vessels, W. C. Price and 
A. R. Philpotts, AC 4031/IB./FP.167, (WA-656-15). 

359. Stresses in Shell, D. N. de G. Allen, AC 4032/Gn.253, 
(WA-708-4), May .1943. 

360. Abstracts of N.D.R.C. Reports Nos. A-91 and A-161, AC 
4144/Gn.259, (WA-708-7), June 1943. 

361. Stresses in H.E. Shell: Comparison of Approximate and 
Relaxation Methods — III, D. G. Sopwith, AC 4252/Gn. 
262, (WA-746-7), June 1943. 

362. Stresses in H.E. Shell: Effect of Driving Band Pressure — 
II, D. G. Sopwith, AC 4254/Gn.263, (WA-746-8), June 
1943. 

363. Stresses in H.E. Shell: Firing Trials in 2 Pr. H.E. Shells 
with Reduced Factors of Safety, D. G. Sopwith, AC 4412/ 
Gn.275, (WA-791-6), July 1943. 

364. Interim Report on the Thermal Decomposition and Burn¬ 
ing of S.C. Cordite at Ordinary Pressure, C. E. H. Bawn, 
AC 4658/PRP.35, (WA-940-10), Aug. 24, 1943. 

365. The Examination of Eroded Metal Surfaces by Means of 
Replica Films, J. F. Allen and P. L. Willmore, AC 5341/ 
Gn.317, (WA-1490-16), December 1943. 

366. Stresses in Guns, D. N. de G. Allen, AC 5342/Gn.318, 
(WA-1490-17), December 1943. 

367. Review of Work Carried out in 1943 in the Armament Re¬ 
search Department, Metallurgical Branch, AC 5467/Met. 
184, (WA-1484-8), January 1944. 

368. Determination of the Coefficients of Friction of Steel on 
Steel at High Velocities, G. Grotsch and E. Plake, AC 
5912/Gn.340 Ball. 161, (RD Translation No. 898), 
(WA-3135-2A), March 1944. (See Item 493.) 

369. Stresses in Thin-Walled Shells, W. E. A. Acum and 
D. N. de G. Allen, AC 5997/Gn.344, (WA-2057-14), 
March 1944. 

370. Note on Longitudinal Strain in Gun Tubes, J. W. Graggs 
and C. J. Tranter, AC 6930/Gn.376, (WA-2961-12), 
September 1944. 

371. A Review of Some Recent Canadian Work on Gun Pro¬ 
pellants, K. J. Laidler, AC 6867/Bal.200, (WA-3005-5), 
September 1944. 


CONFIDENTIAL 





646 


BIBLIOGRAPHY 


372. The Design of Driving Bands, R. Beeching, AC 7199/ 
Gn.389 (ADD Tech. Report 32/44), (WA-3145-6), 
September 1944. 

373. Stresses in Thin-Walled Shells, W. E. A. Acum and D. N. 
de G. Allen, AC 7586/Gn.401, (WA-3822-14), De¬ 
cember 1944. 

374. Tabulation of the Hunt and Hinds System of Internal 
Ballistics, M. M. Nicolsen, AC 8335/Ball.281, (WA- 
5173-7), June 1945. 

375. The Choice of Calibre for Bomber and Fighter Aircraft, 
RD Armament Report 1, (WA-3208-la), September 
1944. 

376. The Firing Stresses in the Q.F. 17-Pr. Gun, G. H. Weston 
and J. B. Goode, ARD Ballistics Report 7/45, (WA- 
3830-6), January 1945. 

377. The Firing Stresses in the Q.F. 2-Pr. Mark 10 Gun 
K. V. Hyde and J. B. Goode, ARD Ballistics Report 
9/45, (WA-3830-7), January 1945. 

378. The Firing Stresses in the Q.F. 6-Pr. 6Cwt. Mk. 3, G. H. 
Weston and J. B. Goode, ARD Ballistics Report 11/45, 
(WA-4104-1), February 1945. 

379. Notes on the Salvage and Recovery of Overmachined Steel 
Parts by Electrodeposition, ARD Electrodeposition 
Memorandum No. 1, (WA-3767-4), January 1944. 

380. Notes on the Inspection of Heavy Electrodeposits of Nickel 
and Chromium , ARD Electrodeposition Memorandum 
No. 5, (WA-3767-4C), May 1943. 

381. Report on Exudation and Cook-Off of Cordite Loaded 
Ammunition in Browning M.G., ARD Explosives Re¬ 
port 111/42, (II-5-3425), May 1942. 

382. Repair of Worn or Over-Machined Steel Components by 
Electrodeposition of Nickel, RD Met. Report 205/41, 
(II-5-1686), April 1941. 

383. Plant for Repair of Gun Barrels of Chambers by Electro¬ 
deposition of Nickel, RD Met. Report 458/41, (II-5- 
1830), September 1941. 

384. Improvements in the Electrodeposition of Chromium: Pro¬ 
duction of Machinable Chromium Deposits, G. E. Gar- 
dam, RD Met. Report 465/41, (WA-776-7c), Septem¬ 
ber 1941. (See Item 471.) 

385. Electrodeposition of Chromium in Gun Barrels; Summary 
of Present Position of Knowledge of the Control of the 
Hardness and Structure of Electrodeposifed Chromium, 
G. E. Gardam, RD Met. Report 400/42, (WA-171-la), 
July 1942. 

386. Notes on the Inspection of Heavy Electrodeposits of Nickel 
and Chromium, ARD Met. Report 235/43, (WA-1258- 
14), October 1943. 

387. German 10.5 “Sabot” Type H.E. Shell, ARD Met. Re¬ 
port 26/44, (WA-1630-8), February 1944. 

388. Compression Properties of Driving Band Copper, ARD 
Met. Report 90/44 (AC 6257/Gn.354) (WA-2136-12), 
April 1944. 


389. Method of Chromium Plating 0.5 inch Browning Barrels: 
Preliminary Process Specification, ARD Met. Report 
159/44, July 1944 [II-5-6092(s)]; (WA-4283-12), Re¬ 
vised April 1945. 

390. Firing Trials with .5 inch Browning M/G Barrels with 
Chromium Plated Bores, ARD Met. Report 206/44, 
(WA-2926-7H), September 1944. 

391. Examination of Chromium Plated German Machine Gun 
and Breech Mechanism, ARD Met. Report 221/44, 
(WA-3054 and WA-3524-1), October 1944 and Adden¬ 
dum, November 1944. 

392. Examination of a 7.5-cm Sabot Type A.P. Projectile, 
ARD Met. Report 253/44, (WA-3650-4), November 
1944. 

393. A Spectroscopic Investigation of the Light Emitted ( 1) in 
the Burning of Calcium Free Cordite and {2) in Secondary 
Gun Flashes Obtained with the Cordite, W. C. Price and 
H. R. Philpotts, A. C. 3870. 

394. The 6 Pdr. Sabot Projectile, G. S. Sanders, ARD Tech. 
Report 2/45, (WA-4153-2), January 1945. 

395. Notes on the Principles and Present Position of Hyper- 
Velocity Guns and Projectiles, G. O. C. Probert, ARD 
Terminal Ballistic Report 22/44, (WA-2905-5A), July 

1944. 

396. Theories of the Burning of Colloidal Propellants, J. Cor¬ 
ner, ARD Theoretical Research Report 2/43, (WA- 
1297-5), October 1943. 

397. Some Theoretical Considerations of the Problem of Gun 
Erosion, C. K. Thornhill, ARD Theoretical Research 
Report 7/43, (WA-1424-2), December 1943. 

398. The Heating of a Gun Barrel by the Propellant Gases, II. 
Comparison between Theoretical and Experimental Values 
of Heat Transfer in Small Arms Weapons, C. K. Thorn¬ 
hill and H. N. Ware, ARD Theoretical Research Report 
29/44 (AC 7091/Gn.384/Ball.-209), (WA-3005-11), 
July 1944. (See Item 403 for revision). 

399. The Stresses in a Film of Fluid Lubricant on the Bore 
Surface of a Gun Barrel during Shot Travel, E. P. Hicks, 
ARD Theoretical Research Report 48/44, (WA-3695-2), 
December 1944. 

400. Report on a Visit to the United States, February-April 
19^5, in Connection with Work on Gun Erosion, ARD 
Theoretical Research Report 14/45, (WA-5071-3), June 

1945. 

401. An Estimate of the Heating of a 3-in. Gun during Firing 
at 90 R.P.M., C. K. Thornhill and E. P. Hicks, ARD 
Theoretical Research Memorandum 15/45, May 1945. 

402. The Pressure of Air and Gas between the Bands of a Lit¬ 
tlejohn Projectile, J. Corner, ARD Theoretical Research 
Memorandum 20/44, (AC 6932/Gn.377) (WA-2870-11), 
August 1944. 

403. Notes on the Continuous Heating of a Gun Barrel during 
Successive Firings, with Revision of Experimental Results 
Derived in ARD Theoretical Research Report No. 29 /44, 
ARD Theoretical Research Memorandum 3/45, [WA- 
2792-2a(l)], January 1945. (See Item 398.) 


CONFIDENTIAL 




BIBLIOGRAPHY 


647 


404. The Life of Gun Barrels, ARD Theoretical Research 
Translation 2/44, (WA-3065-9), September 1944. 

405. Barrel Wear of a .5-in. Cal. Browning Machine Gun for 
Aircraft; a Review of the Principal Proceedings of the 
Ordnance Board, 1940-March 1944, A. F. Burstall, ADD 
Technical Report 5/44, (I-A-243), July 1944. 

406. The Life of the .50" Browning Barrel, C. F. Austin and 
R. Beeching, ADD Technical Report 10/45 (First 
Report of the Barrel Life Panel of the Armament Design 
and Research Departments), April 1945 (WA-4517-11). 
Addendum, B. C. Brookes, Report 10a/45, (WA-5635- 
4), December 1945. 

407. Barrel Wear of .5-in. Cal. Browning Machine Gun for 
Aircraft: A Review of Items of Special Interest to the Panel 
on Barrel Life of Machine Guns Taken from the Recent 
Reports of Division 1 of the National Defense Research 
Committee ( N.D.R.C.), U.S.A., A. F. Burstall, ADD 
Technical Report 12/44, (WA-2177-7b), March 1944. 

408. Summary from Geophysical Laboratory ( U.S.A .) Report, 
May 1944 N.D.R.C. Div. 1 — 320; Chromium Plated and 
Nitrided Liners and 9% lb. Barrels, Barrel Life Panel 
B.R. 561, (WA-3506-ld), July 1944. 

409. Summary of American Work (N.D.R.C. Div. 1) Relevant 
to Erosion and .50-in. Browning Barrel Life; January- 
March, part April and May, 1944, C. F. Austin, Barrel 
Life Panel B.R. 578, (WA-3506-lb), August 1944. (Con¬ 
tinuation of ADD Tech. Report 12/44, Item 407.) 

410. Summary of American Work (N.D.R.C. Div. 1) Relevant 
to Erosion and .50 in. Browning Barrel Life, March—June 
30th, 1944, C. F. Austin, Barrel Life Panel B.R. 659, 
(WA-3506-lc), September 1944. 

411. Note on the Cooling of .500" Browning Barrels, Barrel 
Life Panel S.A. 2542, (II-5-6319), March 1944. 

412. Results of Trials on .5-in. Cal. Browning Barrels, Tem¬ 
perature Indicating Paints and their Use in Barrel Rejec¬ 
tion, Barrel Life Panel S.A. 2542/38R, [WA-2534-llg 
(1) ], July 1944. 

413. Results of Trial on .5-in. Cal. Browning Barrels. Accu¬ 
racy, Velocity and Bore Wear when Firing Various Cycles 
of A.P.M2 through Barrels Having Stellite Liners, Barrel 
Life Panel S.A. 2542/82R, (WA-3750-5B), December 
1944. 

414. .50" Barrels Nitrided and Chromium Plated in USA, 
Barrel Life Panel S.A. 2542/108R, [II-5-7025(s)], 
April 1945. 

415. .50" Cadmium Plated Barrels [Bullets], Barrel Life Panel 
S.A. 2542/112R, (WA-4545-3M), May 1945. 

416. .50" Chromium Plated Barrels from U. S.A. ( B.A.C. Lot 
1), Barrel Life Panel S.A. 2542/115R, (WA-5432-2), 
June 1945. 

417. .50" Barrels Nitrided and Chromium Plated in U.S.A. 
(B.A.C. Lot 2), Barrel Life Panel S.A. 2542/119R, (WA- 
5432-6), August 1945. 

418. Nitrided and Chromium Plated Barrels, Barrel Life Panel 
S.A. 2542/127R, (WA-5658-8), November 1945. 


419. 0.50-inch Browning Chrqmium Plated Barrel Life, Incor¬ 
porating the Effect of Cooling by Air Flow, with the Use 
of Cooling Ducts Designed by N.A.F.D.U., Pendine 
Report 112/^5, (WA-5164-5), July 1945. 

420. 0.50-inch Browning Barrel Life, Incorporating the Effect 
of Cooling by Air Flow, with the Use of Ducts Designed by 
N.A.F.D.U., Pendine Report 290/44, [II-5-7017(s)], 
April 1945. 

421. The Nitrogen Content and Constitution of the Hard “White 
Layer” in Eroded Guns, Brynmor Jones, University Col¬ 
lege, Cardiff, Research Dept. (Woolwich), Extra-Mural 
Report No. 8, (II-5-2042), 1942. 

422. The Engravement of Shell on Firing and its Possible Rela¬ 
tion to Wear in Guns, J. F. Allen, S.R.I. (FRG) Reports, 
(WA-2877-3A), (AC 6971/Gn.381 Ball. 202, October 
1944). 

423. Chromium Plating of Bores of 0.5-inch Browning Barrels, 
Informal Communication from A. W. Hothersall (Arma¬ 
ment Research Department) to British Central Scienti¬ 
fic Office, (II-5-6295), October 1944. 

424. Stresses and Strains in Thick and Thin Tubes with Semi- 
Infinite Pressure Distribution, D. G. Sopwith, Eng. Div. 
Report 116/45 (AC 8130/Gn.428), (WA-4334-12), 
March 1945. 

REPORTS ISSUED BY OTHER BRITISH 
GOVERNMENT OFFICES h 

425. High-Velocity Guns on the Gerlich and Smoothbore Princi¬ 
ples, Ordnance Board (Gr. Brit.), Proceeding No. 
11379, Mar. 13, 1941. 

426. German 28/20-mm A/T Gun (Model 41): Pressures in 
Petrol Tanks Due to the Impact of Projectiles, Ordnance 
Board (Gr. Brit.), Proceeding No. 15,750, (W-137-47), 
Jan. 9,1942. 

427. Q. F. 3.7-inch gun Mark VI: (i) Mechanical Break-down 
to be Expected in Prolonged Firing; (ii) Possibility of 
“Cook-off", Ordnance Board (Gr. Brit.), Proceeding 
No. 27,604, (WA-2140-23), May 12, 1944. 

428. 20-mm Hispano: “Cook-off" in Hot Guns: (i) Results of 
Trials; (ii) Explosions in Lips of Belt Feed Mechanism, 
Ordnance Board (Gr. Brit.), Proceeding No. 30,417, 
(WA-3980-12), Feb. 26, 1945. 

429. “Effect of Increased Temperature on Firing", B. L. 16- 
inch Mark IV Gun for Naval Service, Ordnance Board 
(Gr. Brit.), Proceeding No. Q 3617, Appendix, Sec. 30, 
(WA-5099-4). 

430. Cooling of .50" Browning Guns in Naval Aircraft, Pre¬ 
liminary Report, Reports Nos. 2, 4 to 8 inch, NAFDU/ 
40/39 Arm.; Report No. 3, NAFDU/40/73 Arm., [II-5- 
6431 (s) to II-5-6438(s), inch], August-October 1944. 
(See Item 302 for summary of these reports.) 


h Item 544 in Additional References of this Bibliography per¬ 
tains to this section. 


CONFIDENTIAL 






648 


BIBLIOGRAPHY 


431. A System of Internal Ballistics, F. R. W. Hunt and G. H. 
Hinds, revised by C. J. Tranter, Military College of 
Science Publication, G. M. II, H. M. Stationery Office, 
(WA-45154J), 1941. 

432. Temperature Indicating Materials, Post Office Engineer¬ 
ing Department, Radio Report No. 1298, (WA-4566-7), 
February 1945. 

433. Some Notes on the Calibration of Thermindex Paints, In¬ 
formal report submitted with letter January 10, 1945 
from L. C. Tyte to G. Comenetz, Geophysical Labora¬ 
tory, Carnegie Institution of Washington (WA-3679- 
6A,B). 

JOURNAL ARTICLES (UNCLASSIFIEDY 

434. “Supersonic Flow over an Inclined Body of Revolution,” 
H. S. Tsien, J. Aeronaut. Sci., Vol. 12, 1938, p. 480. 

435. “A Contribution to the Thermodynamical Theory of 
Explosions,” J. B. Henderson, H. P. Hasse, Proc. Roy. 
Soc. (London), Vol. A100, 1922, p. 461. 

436. Proudman, Proc. Roy. Soc. 100A, 289 (1922). 

437. “The Air Pressure on a Cone Moving at High Speeds,” 
G. I. Taylor, J. W. Maccoll, Proc. Roy. Soc. (London), 
Vol. A139, 1933, p. 278. 

438. “The Conical Shock Wave Formed by a Cone Moving 
at a High Speed,” J. W. Maccoll, Proc. Roy. Soc. (Lon¬ 
don), Vol. A159, 1937, p. 459. 

439. Mansell, Trans. Roy. Soc. (London), 207A, 243 (1907). 

440. “On the Equation of State of Propellant Gases,” A. D. 
Crow, W. E. Grimshaw, Trans. Roy. Soc. (London), 
Vol. A230, 1931, p. 39. 

441. “Combustion of Colloidal Propellants,” A. D. Crow, 
W. E. Grimshaw, Trans. Roy. Soc. (London), Vol. A230, 
1931, p. 387. 

442. “Researches on Explosives,” A. Noble, F. A. Abel, 

Trans. Roy. Soc. (London), Vol. A -, 1875, p. 122. 

Reprinted in Artillery and Explosives, 1906, p. 99. 

443. “Researches on Explosives,” A. Noble, Trans. Roy. Soc. 
(London), Vol. A205, 1905, p. 201. 

444. “The Aerodynamics of a Spinning Shell,” R. H. Fowler, 
E. G. Gallop, G. N. H. Lock, H. W. Richmond, Trans. 
Roy. Soc. (London), Vol. A221, 1920, p. 296. 

445. “Campo Aerodinamico a Velocita Iperacustica Attorno 
a un Solido di Rivoluzione a Prora Acuminata,” C. 
Ferrari, UAerotecnica, Vol. 16, No. 2, 1936, p. 121. 

446. “Campi di Corrente Impersonora Attorno a Solidi di 
Rivoluzione,” C. Ferrari, VAerotecnica, Vol. 17, No. 6, 
1937, p. 507. 

447. “Valence Relations Among the Metal Carbonyls,” A. A. 
Blanchard, Chem. Rev., Vol. 26, No. 3, June 1940, p. 409. 


•The following items in Additional References of this Bib¬ 
liography pertain to items 534, 536, 537, 542, and 559 of this 
section. 


448. “The Metal Carbonyls,” A. A. Blanchard, Science, Vol. 
94, No. 2440, Oct. 3, 1941, p. 311. 

449. “Die Bildung von Eisen-, Kobalt- und Nickelcarbonyl 
durch Hochdrucksnythese aus Halogeniden in ver- 
gleichen der Darstellung,” H. Hieber, H. Behrens, U. 
Teller, A. anorg. allgem. Chem., Vol. 249, 1942, p. 26. 

450. “Arbeitsbilanz beim Schuss aus einem Gewehr,” C. 
Cranz, R. Rothe, Z. ges. Schiess-u. Sprengstoffw., Vol. 3, 
1908, pp. 301, 327, 474. 

451. “Decomposition, Heat of Combustion and Temperature 
of Explosion of Explosives,” O. Poppenberg and E. 
Stephan, Z. ges. Schiess-u. Sprengstoffw., Vol. 4, 1909, 

p. 281. 

452. “Decomposition of Powder in Guns during Firing,” O. 
Poppenberg, E. Stephan, Z. ges. Schiess-u. Sprengstoffw., 
Vol. 4, 1909, p. 388. 

453. “Beitrage zur thermodynamischen Behandlung ex- 
plosibler Vorgange, I,” A. Schmidt, Z. ges. Schiess-u. 
Sprengstoffw., Vol. 24, No. 2, 1929, p. 41. 

454. “Intermolecular Forces and the Properties of Gases,” 
J. O. Hirschfelder, W. E. Rosenveare, J. Phys. Chem., 
Vol. 43, 1939, p. 15. 

455. J. O. Hirschfelder, R. B. Ewell, J. R. Roebuck, J. Chem. 
Phys., Vol. 6, 1938, p. 205. 

456. J. O. Hirschfelder, F. T. McClure, I. F. Weeks, J. Chem. 
Phys., Vol. 10, 1942, p. 201. 

457. “Decomposition of Nitrocellulose,” R. Robertson, S. S. 
Napier, J. Chem. Soc., 1907, p. 761. 

458. “Decomposition of Nitroglycerine,” R. Robertson, J . 
Chem. Soc., 1909, p. 1241. 

459. “The Deposition of Chromium from Solutions of 
Chromic and Chromous Salts,” Charles Kasper, J. Re¬ 
search Natl. Bur. Standards, Vol. 11, 1933, p. 515. 

460. “Mechanism of Chromium Deposition from the Chrom¬ 
ic Acid Bath,” Charles Kasper, J. Research Natl. Bur. 
Standards, Vol. 14, 1935, p. 700. 

461. “Cathodic Deposition of Chromium Nickel Alloy,” 
M. F. Skalozubov, A. S. Vlasova, Chem. Abs., Vol. 35, 
1941, p. 1323. 

462. “Cathodic Deposition of Iron-Chromium-Nickel Al¬ 
loy,” M. F. Skalozubov, I. A. Gencharova, Chem. Abs., 
Vol. 35, 1941, p. 1323. 

463. “The Electrodeposition of Tungsten from Aqueous 
Solutions,” C. G. Fink, F. L. Jones, Trans. Electrochem. 
Soc., Vol. 59, 1931, p. 461. 

464. “The Co-deposition of Tungsten and Iron from Aqueous 
Solutions,” M. L. Holt, Trans. Electrochem. Soc., Vol. 
66, 1934, p. 453. 

465. “Metals Co-deposited with Tungsten from the Alkaline 
Tungsten Plating Bath,” M. L. Holt, Trans. Electro¬ 
chem. Soc., Vol. 71, 1937, p. 301. 

466. “Electrodeposition of Nickel-Tungsten Alloys from an 
Acid Plating Bath,” M. L. Holt, M. L. Nielsen, Trans. 
Electrochem. Soc., Vol. 82, 1942, p. 193. 


CONFIDENTIAL 




BIBLIOGRAPHY 


649 


467. “Electrodeposition of Iron-Tungsten Alloys from an 
Acid Plating Bath,” M. L. Holt, R. E. Black, Trans. 
Electrochem. Soc., Vol. 82, 1942, p. 205. 

468. “Electrodeposition of Cobalt-Tungsten Alloys from an 
Acid Plating Bath,” M. L. Holt, R. E. Black, P. F. 
Hoglund, Trans. Electrochem. Soc., Vol. 84, 1943, p. 353. 

469. “The Electrowinning of Chromium from Trivalent Salt 
Solutions,” R. R. Lloyd, W. T. Rawles, R. G. Feeney, 
Trans. Electrochem. Soc., Vol. 89, (Preprint No. 20), 
April 1946. (See Item 522.) 

470. “Electrolytic Polishing of Stainless Steel and Other 
Metals,” O. Zmeskal, Metal Finishing, Vol. 43, 1945, 

p. 280. 

471. “The Production of Machineable Chromium Deposits,” 
G. E. Gardam, J. Electrodepositors ’ Tech. Soc., Vol. 20, 
1945, p. 69. 

472. “Electrolytic Deposition of Alloys of Tungsten, Nickel, 
and Copper from Water Solutions,” L. N. Goltz, Z. N. 
Kahrlamov, J. Applied Chem. ( U.S.S.R. ), Vol. 9, 1936, 
p. 640. 

473. “Electrolytic Deposition of Tungsten and its Practical 
Utilization,” S. I. Sklyarenko, O. S. Druzhinia, M. M. 
Masai’tseva, J. Applied Chem. (U.S.S.R.), Vol. 13, 
1940, p. 1926. 

474. “Eloctroplating with Tungsten and Molybdenum,” P. P. 
Belyaev, A. I. Lipovetskaya, Korroziya i Borba s Nei, 
Vol. 6, No. 247, 1940. 

475. “The Erosion of Gun Tubes and Heat Phenomena in 
the Bore of a Gun,” H. J. Jones, Engineer, Vol. Ill, 
1911, p. 294. 

476. “The Erosion of Guns,” H. M. Howe, Trans. Am. Inst. 
Min. Met. Engrs., Vol. 58, 1918, p. 513. 

477. “Nitrogen in Steel and the Erosion of Guns,” H. E. 
Wheeler, Trans. Am. Inst. Min. Met. Engrs., Vol. 67, 
1922, p. 257. 

478. “Some Notes on the Effect of Nitrogen on Steel,” O. A. 
Knight, H. B. Northrup, Chem. & Met. Eng. Vol. 23, 
1920, p. 1107. 

479. “The Erosion of Guns,” R. H. Greaves, H. H. Abram, 
S. H. Rees, J. Iron and Steel Inst., Vol. 119, 1929, p. 113. 

480. “La Temperature de Combustion de la Poudre sans 
Fumee en Fonction de sa Composition Chimique,” 
Part I, A. L. Th. Moesveld; Parts II and III, G. de 
Bruin, P. F. de Pauw * Comm, de la Ste. Anme. Fabrique 
Neerlandaises d’Explosifs, Nos. 7, 8, 9, 1928, 1929. 

481. “Carburization as a Factor in the Erosion of Machine 
Gun Barrels,” W. W. deSveshnikoff, Army Ordnance, 
Vol. 5, 1925, p. 794. 

482. “[White Layer] Symposium,” Army Ordnance, Vol. 5, 
1925, p. 797. (This symposium included the following 
five papers: “The Cause of the White Layer,” Henry 
Fay; “Erosion of Machine Gun Barrels,” W. T. Gorton; 
“Additional Tests Needed to Determine Value of Var¬ 
ious Steels,” J. S. Vanick; “The Nitrogen Theory of 
Erosion,” H. E. Wheeler; “Old and New Theories in 
Gun Erosion,” A. G. Zimmermann.) 


483. “The White Layer in Gun Tubes and its Relation to the 
Case of Nitrided Chromium-Aluminum Steel,” H. H. 
Lester, Trans. Am. Soc. Steel Treating, Vol. 16, No. 5 
1929. 

484. “The Sulfides of Zinc, Cadmium, and Mercury; Their 
Crystalline Forms and Genetic Conditions,” E. T. Allen, 
J. L. Crenshaw, H. E. Merwin, Am. J. Sci., Vol. 34, 
1912, p. 341. 

485. “Resistance of Slender Bodies Moving with Supersonic 
Velocities with Special Reference to Projectiles,” T. von 
Karman, N. B. Moore, Trans. Am. Soc. Mech. Engrs., 
Vol. 54, 1932, p. 303. 

486. “Neuere Ergebnissee der Funkenkinematographie,” H. 
Schardin, W. Struth, Z. Tech. Physik, Vol. 18, No. 11, 
November 1937. Translation: “Recent Results in Spark 
Cinematography,” Translation No. 107, David W. 
Taylor Model Basin, Navy Dept., December 1942. 

487. “Combustion of Explosives in Closed Vessels. Compari¬ 
son of Experimental and Calculated Temperature and 
Pressure,” H. Muraour, G. Aunis, Chaleur et Ind., Vol. 
20, 1939, p. 31. 

488. “Notes de balistique interieure,” H. Muraour, Mem. 
artillerie frangaise, Vol. 4, 1925, p. 455. 

489. “Pre-engraved Projectiles,” (“Les obus rayes”) P. 
Charbonnier, Mem. artillerie frangaise, Vol. 6, 1927, p. 
3. 

490. “The Wear of Rifle and Automatic Arms Barrels and 
their Metallographic Analysis,” (“Usure des canons 
d’armes portatives et automatiques, et leur analyse 
metallographique”), P. Felsztyn, S. Spiewak, Mem. 
artillerie frangaise, Vol. 17, 1938, p. 283. 

491. “Resistance de forcement des projectiles dans le canon 
et son etude par le calcul,” M. C. Spetzler, Mem. 
artillerie frangaise, Vol. 17, No. 68, 1938, p. 869. 

492. “Ein Verfahren zur Messung schnellveranderlicher 
Oberflachentemperaturen und seine Anwendung in 
Schusswaffenlaufen,” P. Hackemann, Luftfahrtfor- 
schungsanstalt Hermann Goring, (E. V. Braunschweig), 
Jan. 27, 1941. (Translation: “A Method for Measure¬ 
ment of Fast Changing Surface Temperature and its 
Application for Small Arms Barrels,” Ordnance Tech¬ 
nical Intelligence Branch, OTIB No. 1243.) 

493. “Bestimmung des Reibungskoeffizienten bei hohen 
Geschwindigkeiten fur Stahl auf Stahl,” G. Grotsch, E. 
Plake, Jahrbuch 1938 der deutschen Luftfahrt-Forsch, 
p. 345. (See Item 368.) 

494. R. Sauer, Luftfahrt-Forsch., Vol. 19, 1942, p. 148. 

495. Busemann, Luftfahrt-Forsch., Vol. 19, 1942, p. 137. 

496. “Flow in a Smooth Straight Pipe at Velocities above and 
below the Velocity of Sound,” W. Frossel, Forschung auf 
dem Gebiete des Ingenieurwesens, Vol. 7, March-April 
1936. (Translation: Tech. Memo. No. 844, National 
Advisory Council for Aeronautics, 1938.) 

497. H. Muraour, Bull. Soc. Chim. Mem. Vol. 41, 1927, p. 
1451. 


CONFIDENTIAL 





650 


BIBLIOGRAPHY 


498. H. Langweiler, Z. Physik, Vol. 19, 1938, p. 416. 

499. A. Mittasch, Z. angew Chem., Vol. 41, 1928, p. 827. 

500. R. H. Kent, Physics, Vol. 7, 1936, p. 319. 

501. J. E. Lennard-Jones, Physica , Vol. 4, 1937, p. 941. 

502. “Apparatus for Optical Studies at High Pressure,” T. 
Poulter, Phys. Rev. Vol. 40, 1932, p. 860. 

BOOKS ( UNCLASSIFIED)> 

503. Metals Handbook, American Society for Metals; Cleve¬ 
land, Ohio, 1939, pp. 368, 390. 

504. Handbook of Chemistry and Physics , 25th ed,; Chemical 
Rubber Publishing Co., Cleveland, Ohio, 1941. 

505. Lehrbuch der Ballistik, C. Cranz, Edwards Brothers, 
Ann Arbor, Mich., 1943. (Cf. Item 144). 

506. Statistical Thermodynamics , R. H. Fowler, E. A. Guggen¬ 
heim, Cambridge University Press, 1939, Chapter 7. 

507. Modern Developments in Fluid Dynamics, Vol. II, S. 
Goldstein, Oxford-Clarendon Press, 1938. 

508. Kinematics of Machines, George L. Guillet, 4th ed., 
John Wiley and Sons, 1940. 

509. Elements of Ordnance, T. J. Hayes, John Wiley and 
Sons, 1938. 

510. Automatic Arms: Their History, Development and Use, 
M. M. Johnson, Jr., C. T. Haven, Morrow and Co., 
1941. 

511. Thermodynamics and Free Energy, G. N. Lewis, M. 
Randall, McGraw-Hill Book Co., 1923. 

512. A Treatise on the Mathematical Theory of Elasticity, 
A. E. H. Love, 4th ed., Cambridge University Press, 
1934, pp. 274-277. 

513. Heat Transmissions, W. H. McAdams, 2nd ed., Mc¬ 
Graw-Hill Book Co., 1942. 

514. Plasticity, A. Nadai, McGraw-Hill Book Co., 1931, 
Chapter 13. 

515. Naval Ordnance (A textbook prepared for the use of 
midshipmen at the U. S. Naval Academy), Officers of 
the U. S. Navy, Annapolis, U. S. Naval Institute, 1939. 

516. Principes de Thermodynamique, Saint Robert, Turin, 
1870, p. 251 ff. 

517. Friction, Stanton, London, 1923, p. 149. 

518. Theory of Elasticity, S. Timoshenko, McGraw-Hill Book 
Co. 

519. Theory of Plates and Shells, S. Timoshenko, McGraw- 
Hill Book Co., 1940. 

520. Text-Book of Ordnance and Gunnery, W. H. Tschappat, 
1st ed., John Wiley and Sons, 1917. 

521. Chemical Computations and Errors, Crumplet and Yoe, 
p. 232. 


j Item 535 in Additional References of this Bibliography per¬ 
tains to this section. 


MISCELLANEOUS DOCUMENTS 

522. The Electrowinning of Chromium, R. S. Dean, Annual 
Report of Metallurgical Division, Fiscal Year 1941, 
U. S. Bureau of Mines Report Investigations 3600,1941. 
(See Item 469.) 

523. A Digest of Ballistic Data on Large Naval Guns Obtained 
for the Bureau of Ordnance, Navy Department, National 
Bureau of Standards, October 1924. 

524. Chromium-Base Alloys, R. M. Parke, F. P. Bens, Cli¬ 
max Molybdenum Company, Mar. 15, 1945. (Report 
submitted to War and Navy Departments for clearance 
prior to publication.) 

525. Performance and Design Criteria for .50 Caliber Gun 
Barrel Coolers for Aircraft, J. A. Broadston, Report No. 
NA-8568, North American Aviation, Inc., 1945. 

526. Report on Gun Barrel Cooling Tests for Mustang Aero¬ 
planes, J. A. Broadston, North American Aviation, Inc. 

527. [Fluoride Baths for Electroplating], Armstrong, Menefee, 
U. S. Patents: 2,145,241; 2,145,745; 2,145,746; 2,160,- 
321; 2,160,322. 

528. Machine for Rifling Guns, V. Bush, U. S. Patent No. 
2,319,206, Jan. 23, 1942. 

529. Improvements in the Manufacture and Production of Col¬ 
oured Coatings Capable of Indicating Temperatures, 
Patent Specification (British) 478,140, Jan. 10, 1938. 

530. Thermetric Colours, Dyestuffs Division, Imperial Chem¬ 
ical Industries, Ltd. 

531. Temperature Indicating Paints, F. G. Nicholls, Austra¬ 
lian and New Zealand Scientific Research Liaison 
(London) Report No. 156, (II-5-2397), August 1942. 

532. The Friction of the Driving Bands of Shells, Part II, 
Lubrication and Friction Report No. 34, Council for 
Scientific and Industrial Research, Australia, (A.C. 
5847/Gn.337, March 1944), (WA-1977-15), November 
1943. 

533. The Mathematical Analysis of dp/dt-p-t Curves Ob¬ 
tained from Closed Vessel Firings, N. S. Mendelshon, 
IBRL Report No. 5, Internal Ballistics Research Lab¬ 
oratory, Valcartier, Quebec. 

ADDITIONAL REFERENCES 

534. G. Jaeger, Wiener Sitzber, Vol. 105, 1896, p. 15. 

535. Gastheorie, II, L. Boltzmann, 1898, sec. 51-61. 

536. J. D. van der Waals, Jr., Versl. K. Akad. Amst, Vol. 10, 
1902, p. 640. 

537. H. Happell, Ann. d. Phys., Vol. 21, 1906, p. 342. 

538. Calorific Values of Smokeless Powders as Affected by 
Variations in Composition, Granulation, etc., C. G. 
Dunkle and N. T. Volsk, Picatinny Arsenal Tech. 
Reports Nos. 231, 620, June 1932, April 1935. 

539. Thermochemical Examination of a Number of Commer¬ 
cial and Experimental Propellants, J. F. Kincaid, OSRD 
1578, Division 8, NDRC. 


CONFIDENTIAL 




BIBLIOGRAPHY 


651 


540. [Tests to Determine Ballistic Characteristics of Projectile, 
A.P., 57/40-mm, J&L A-1944], Firing Records Nos. 
P-35551, P-35549, Ordnance Program No. 5829, APG, 
July 5, 10, 11, 1945. 

541. P. R. Kosting and R. E. Peterson, WA Report WAL 
731/59. 

542. “Some Two-Dimensional Problems in Conduction of 
Heat with Circular Symmetry,” H. S. Carslaw and J. C. 
Jaeger, Proc. Bond. Math. Soc., Vol. 46, 1939-40, p. 361. 

543. Test of 76-mm Sabot-Projectile, Firing Record No. 
P-35485, APG, Mar. 14, 1945. 

544. Ordnance Board (Gr. Brit.), Proceedings No. 30,047. 

545. [ Report on Breech Injection Devices ], Purdue University. 

546. Stresses in Shell with Wall Ratio 1.5, G. C. Evans, 
T.D.B.S. Report No. 20, OCO, January 1944. 

547. Direct Measurement of Burning Rates by an Electric 
Timing Method, B. L. Crawford, Jr., and C. Huggett, 
OSRD-4009, NRDC Report A 286, Univ. of Minnesota, 
August 1945. 

548. The Final Reactions of the Burning of Nitrocellulose, C. 
P. Fennimore, Report No. 464, BRL-APG, Apr. 29, 
1944. 

549. Experiments on Ignition of Nitrocellulose, C. P. Fenni¬ 
more and J. H. Frazer, Report No. 465, BRL-APG, 
May 12,1944. 

550. A Comparison of Antiaircraft Guns of Various Calibers, 
R. H. Kent, Report No. 125, BRL-APG, Dec. 1, 1938. 

551. Means of Obtaining Greater Armor Penetration from Anti¬ 
tank Guns, R. H. Kent, Report No. 214, BRG-APG, 
Jan. 8,1941. 


552. Gun Erosion: Development of Standardized Erosion Tests, 
P. R. Kosting, Experimental Report No. WAL 731/72, 
WA, Nov. 15, 1943. 

553. Tables for Interior Ballistics, A. A. Bennett, Document 
No. 2039, OCO, Washington, D. C., April 1921. 

554. Project for a New Table for Interior Ballistics for Multi- 
perforated Powder, J. P. Vinti, Report No. 402, BRL- 
APG. 

555. Some Comments on Effect of Adding Length to the After¬ 
body of a Square-Based, Conical-Headed Projectile, R. N. 
Thomas, Report No. 543, BRL-APG, Apr. 23, 1945. 

556. An Investigation into the Behaviour of the Shell Driving 
Band from the Aspect of Gun Design, K. V. Hyde and G. 
H. Weston, A.C. 4758/Gn.294 (ARD Ball. Report 
67/43), (WA-1019-4), September 1943. 

557. Gun Tubes—Stresses and Deformations of Circular Cylin¬ 
ders. Part I: Mathematical Analysis for Cylinders Under 
Uniform Pressure, R. Beeuwkes, Jr., and J. H. Laning, 
Jr., Report No. WAL 660/16, WA. 

558. Progress Report on Behavior of Metals Under Dynamic 
Conditions, (NS-109): The Influence of Impact Velocity 
on the Tensile Properties of Some Metals and Alloys, D. 
S. Clark and P. E. Duwez, California Institute of 
Technology, OSRD 3837 (War Metallurgy Division 
M-288), June 19, 1944. 

559. “A 1-Fin Process,” Metals and Alloys, 1945. 

560. Firing Tests to Determine the Armor Plate Penetrating 
Characteristics of Projectile, A.P., 57/40-mm. Dwg. J&L 
A-1944 , Firing Record No. P-35543, Ordnance Program 
No. 5829, APG, June 29-July 4, 1945. 


CONFIDENTIAL 




OSRD APPOINTEES 


division 1 

Chief 

L. H. Adams 

Deputy Chief 
H. B. Allen 


Members 


W. Bleakney 
L. J. Briggs 
R. Eksergian 


C. E. MacQuigg 
E. L. Rose 
E. R. Weidlein 


Secretary 
J. S. Burlew 


Special Assistants 


J. W. GREIG a J. F. ScHAIRER 

J. P. Marble J. A. TenBrook 


Engineers b 

E. Bainbridge W. H. Shallenberger 

V. Wichum 


Technical Aides 


H. L. Black 
J. S. Burlew 
Grace L. Hart 
0. H. Kneen 


L. E. Line, Jr. 

J. P. Marble 
Rita G. Schubert 
N. H. Smith 


Helen M. Watson 


a Title of appointment was “Consultant,” but functions were those of 
a Special Assistant. 

b Because of the emphasis placed upon the engineering phases of the 
Division’s activities at the time of the reorganization of Division 1, on 
October 1, 1944 the personnel was supplemented by engineers, two of 
whom were assigned by the Engineering and Transitions Office, NDRC. 


652 


CONFIDENTIAL 



CONTRACT NUMBERS, CONTRACTORS, AND SUBJECTS OF CONTRACTS 


Contract No. 

Contractor 

Subject 

OEMsr- 51 

Carnegie Institution of Washington, 
Geophysical Laboratory, 
Washington, D. C. 

Investigation of gun erosion and other ordnance 
problems. 

OEMsr- 430 

Western Electric Company, Inc., 

Bell Telephone Laboratories, 
Murray Hill, N. J. 

Electron diffraction studies of altered layers on 
steel and ferrous alloy specimens. 

OEMsr- 463 a 

The Johns Hopkins University, 
Baltimore, Maryland 

Determination of the effect of carbon monoxide 
and other gases on gun steel. 

OEMsr- 465 

Johnson Automatics, Inc., 

Boston, Mass. 

Development of erosion-resistant liners and coat¬ 
ings in guns, and methods of testing such mate¬ 
rials in the form of short liners. 

OEMsr- 467 

Jones and Lamson Machine Co., 
Springfield, Vt. 

Development, design, and construction of rifled 
gun barrels and related equipment. 

OEMsr- 516 b 

The Catholic University of America, 
Washington, D. C. 

Theoretical interior ballistic investigations, and the 
development of experimental methods for deter¬ 
mining stresses in projectiles and in gun tubes. 

OEMsr- 533» 

The Franklin Institute of the State of 
Pennsylvania, 

Philadelphia, Penna. 

Investigation of gun erosion and other ordnance 
problems. 

OEMsr- 534 

Bryant Chucking Grinder Co., 
Springfield, Vt. 

Design, development, and construction of special 
projectiles.® 

OEMsr- 536 

Leeds & Northrup Company, 
Philadelphia, Penna. 

Development of methods and equipment for re¬ 
cording temperature-time measurements in gun 
barrels. 

OEMsr- 537 

President and Fellows of 

Harvard College, 

Cambridge, Mass. 

Metallurgical studies of erosion effects on alloy 
specimens. 

OEMsr- 598 

The Rector and Visitors of the 
University of Virginia, 
Charlottesville, Va. 

Investigation of centrifugal effects in projectiles. 

OEMsr- 608 

Massachusetts Institute of Technology, 
Cambridge, Mass. 

Development of test specimens of erosion-resistant 
metallic coatings by various techniques. 

OEMsr- 613 

Armour Research Foundation, 

Chicago, Ill. 

Performance of supersonic cavitation erosion tests. 

OEMsr- 629 a 

Crane Co. 

Chicago, Ill. 

Design, development, construction, and testing of 
gun barrel liners. 

OEMsr- 668 

The University of New Mexico, 
Albuquerque, N. M. 

Design, development, and construction of sub¬ 
caliber projectiles. 


a Contract continued by Office of the Chief of Ordnance, War Department. 

b Contract continued by Bureau of Ordnance, Navy Department. 

c Beginning January 1, 1944, this work was continued under Contract OEMsr-467. 


CONFIDENTIAL 


653 






654 


CONTRACT NUMBERS, CONTRACTORS, AND SUBJECTS OF CONTRACTS ( Continued ) 


Contract No. 
OEMsr- 733 d 
OEMsr- 746 
OEMsr- 865 

OEMsr- 886 

OEMsr- 915 b 

OEMsr-1038 

OEMsr-1184 

OEMsr-1205 

OEMsr-1273 
OEMsr-1318 
OEMsr-1320 
OEMsr-1330 b 

OEMsr-1368 
OEMsr-1414 
OEMsr-1424 
OEMsr-1433 
OEMsr-1438 


Contractor 


Subject 


Duke University, 

Durham, N. C. 

Johnson Automatics, Inc. 

Boston, Mass. 

General Electric Company, 

Lamp Department, 

Cleveland, Ohio 

Arthur D. Little, Inc. 

Cambridge, Mass. 

Westinghouse Electric Corp., 

Research Laboratories, 

E. Pittsburgh, Penn. 

Duke University, 

Durham, N. C. 

Western Electric Company, Inc., 

Bell Telephone Laboratories, 
Murray Hill, N. J. 

Westinghouse Electric Corporation, 
Lamp Division, 

Bloomfield, N. J. 

Climax Molybdenum Company, 
Detroit, Michigan 

Yale University, 

New Haven, Conn. 

Climax Molybdenum Company, 

New York, N. Y. 

Union Carbide & Carbon Research 
Laboratories, Inc. 

Niagara Falls, N. Y. 

Remington Arms Company, Inc. 

Ilion, N. Y. 

Crane Co., 

Chicago, Ill. 

Industrial Research Laboratories, Ltd. 
Los Angeles, Cal. 

Johnson Automatics, Inc. 

Boston, Mass. 

Remington Arms Company, Inc. 

Ilion, N. Y. 


Studies of the burning rates of propellant powders. 

Development of a belt-fed 20-mm automatic can¬ 
non for aircraft installation. 

Preparation of short liners of erosion-resistant 
materials. 


Development of lightweight materials and methods 
of their fabrication for use in subcaliber projec¬ 
tiles. 

Development of erosion-resistant materials for use 
as gun liners. 

Theoretical investigations of the thermal effects in 
gun erosion. 

Investigation of vapor-phase plating of erosion- 
resistant materials from metal carbonyls. 

Development of molybdenum for use as gun liners; 
investigation of erosion-resistant, electrodepos- 
ited tantalum. 

Development of erosion-resistant materials, and 
preparation of gun liners of such materials. 

Preparation of chromium carbonyl and of carbonyl- 
deposited chromium plates. 

Development of methods for semicommercial prep¬ 
aration of molybdenum and tungsten carbonyls. 

Development of hot-hard alloys suitable for gun 
liners. 


Development of sabot-projectiles. 

Preparation of 2,250 stellite-lined caliber .50 air¬ 
craft machine gun barrels. 

Development of erosion-resistant, centrifugal-cast 
liners for gun tubes. 

Production of machine gun barrels with erosion- 
resistant liners. 

Production of machine gun barrels with erosion- 
resistant liners, including application of draw 
rifling to this purpose. 


b Contract continued by Bureau of Ordnance, Navy Department. 

transferred to Section H, Division 3. 


CONFIDENTIAL 







CONTRACT NUMBERS, CONTRACTORS, AND SUBJECTS OF CONTRACTS ( Continued ) 655 


Contract No. Contractor 


OEMsr-1444 

Chrome Gage Corporation, 
Philadelphia, Penn. 

Development of pilot plant for chromium plating 
caliber .50 machine gun barrels and production 
of not more than 10,000 such barrels. 

OEMsr-1473 

The A. F. Holden Company, 

New Haven, Conn. 

Investigation of the bonding of hot-hard liners in 
gun tubes, and of heat-treating composite 
barrels. 

OEMsr-1494 

Al-Fin Corporation, 

Jamaica, N. Y. 

Application of thermite-welding process to securing 
of erosion-resistant liners in machine gun barrels. 

OEMsr-1499 

Midvale Company, 

Philadelphia, Penn. 

Manufacture of six special 90-mm gun tubes 
(under purchase agreement). 

OEMsr-1499' 

National Bureau of Standards, 
Electrochemical Section, 
Washington, D. C. 

Improvement in plating processes, and their ap¬ 
plication to gun tubes. 

OEMsr-1499' 

National Bureau of Standards, 
Inductance and Capacitance 
Section, 

Washington, D. C. 

Experimental investigation of bore friction and 
other phases of interior ballistics. 


'Several transfers of funds from OSRD; work continued after 10/31/45 by a transfer of funds from the Navy Department. 


CONFIDENTIAL 






SERVICE PROJECT NUMBERS 



The projects listed below were transmitted to the Office of the Executive 

Secretary, OSRD, from the War or Navy Department through either 
the War Department Liaison Officer for NDRC or the Office of Research 
and Inventions (formerly the Coordinator of Research and Develop¬ 
ment), Navy Department. 

Service 
Project No. 

Title 

OD-32 

OD-42 

Army Projects 

Hydrostatic Testing of Shells. 

Investigation of Methods for the Calculation of the Stresses in the Bases of High-Explosive 
Shell. 

OD-52 

OD-52 Ext. 

OD-52 Ext. 

OD-52 Ext. 

Gun Erosion, Including Hypervelocity Gun Studies. 

Stellite Liner for Caliber .60 Machine Gun Barrels. 

Stellite Liner for Caliber .30 Machine Gun Barrels. 

Development of a Satisfactory Production Process for Nitriding and Chromium Plating 
Caliber .50 Aircraft Machine Gun Barrels. 

OD-52 Ext. 

OD-52 Ext. 

OD-154 

Chromium Plating of Caliber .60 T17E3 Machine Gun Barrels. 

Study of the Problem of Excessive Bore Erosion in High Velocity Aircraft Weapons. 3 
Theory and Design of Muzzle Brakes. 15 

NO-21 

NO-23 

NO-23 Ext. 

NO-26 

NO-26 Ext. 

Navy Projects 

Shell Stresses. 

Gun Erosion. 

Development of Erosion-Resistant Short Liners for the Proposed 3-In./70-Cal. Naval Gun. 
Hypervelocity Guns and Projectiles. 

Advisory Service in Regard to the Heating and Cooling Problems Connected with the 3-In./70- 
Cal. Rapid-Fire Gun. 

NO-124 

NO-202 

Development of an Improved 20-mm Aircraft Cannon to Replace the Hispano-Suiza, AN-M2. 
Ballistic Calculations. 


“Division 1 served only in an advisory capacity to the Ordnance Department. 
b This project was transferred to Division 2, NDltC, at an early stage. 


656 


CONFIDENTIAL 








INDEX 


■i 

The subject indexes of all STR volumes are combined in a master index printed in a separate volume. 
For access to the index volume consult the Army or Navy Agency listed on the reverse of the half-title page. 

/ 


Acceleration of projectiles, 79, 133-134 
Accelerometers 

capacitance-resistance type, 90 
crystal, 80, 90 

differentiators for velocimeter, 89-90 
mutual-inductance type, 89-90 
Accuracy-life of caliber .50 gun barrels, 
450-451 

chromium-plated, 458, 464-466, 473, 
478 

stellite-lined, 450-451, 473, 478 
Acetylene in quenched powder gas, 32 
Adaptors for tapered-bore guns 
chromium-plated, 580 
design, 575-576 
for 57/40-mm gun, 571 
tests, 579 

wear after firing, 579 
Adiabatic compression of gases 
apparatus, 231-232, 306-307 
maximum temperature and pressure, 
305 

pressure as function of time, 306 
studies on powder gas reactions with 
steel, 305-307 

Adiabatic flame temperature of pow¬ 
ders, 324-325 

see also Powder gas, temperature 
calculation, 38-39 

correlation with erosiveness of pow¬ 
der, 324 

correlation with heat input, 318-319, 
324 

effect on gun barrel heating, 325 
energy released during burning, 38 
ADP crystal (ammonium diphenyl 
phosphate), 90 

Air blasts for cooling guns, 121, 126- 
127 

Air-cooled guns 
cook-off tests, 121 
effect on performance, 127 
Aircraft machine gun barrels 
caliber .30; 471-472 
caliber .50; 458, 470-471 
chromium-plated, 458-500 
nitrided, 458-472 
stellite-lined, 445-453, 473-484 
Alabandite (gun erosion product), 
236 

Al-Fin Corporation, aluminum-clad 
gun barrels, 480 

Alloys 

chromium, 360-365, 367-369, 406- 
407, 416-417 

cobalt, 370, 376-377, 391-407 


electroplated coatings, 416-417 
hot hard, 406 

see also Stellite, metallurgy and 
properties 
iron, 406-407 
see also Steel 

molybdenum, 374-377, 383-387 
nickel, 351, 376-377, 406-407 
stellite, 391-413, 443-457, 473-484 
TEW, 335, 352, 407 
tungsten, 376, 417, 434 
Alpha iron, definition, 248 
Altered layers on gun bores 

see Bore surface, thermal transfor¬ 
mation; Erosion of guns, bore- 
surface reactions; Erosion prod¬ 
ucts in gun bores 
Aluminum 
gun barrels, 480-481 
thermochemical resistance, 354 
windshields for projectiles, 583 
Alundum crucibles for melting molyb¬ 
denum alloys, 386-387 
Ammonia 

for decoppering, 234-235 
in powder gases, 272, 301-302 
reaction with carbon monoxide, 293 
Ammoniated water, use in detecting 
defective projectile shells, 536 
Ammonium diphenyl phosphate crys¬ 
tal, 90 
AN-M2 gun 
see Johnson 20-mm gun 
Antiaircraft guns, 118-119, 124, 133 
Armor-piercing projectiles, 189-192 
advantages of hypervelocity projec¬ 
tiles, 9-10 
cap, 191 

deformable projectiles, 191 
De Marre formula, 587-588, 626 
muzzle velocity, 192 
nondeformable projectile, 191-192 
projectile material, 192 
sabot-projectiles, 170, 557-558 
scale effect, 189 
skirted projectiles, 585-590 
specific limit energy, 189-191 
steel, 191 

subcaliber, 192, 626-628 
tests with 57/40-mm tapered-bore 
gun, 585-590 

tungsten carbide cored, 190, 192 
Artillery-type projectiles for erosion¬ 
testing gun 

comparison with pre-engraved pro¬ 
jectiles, 595 


erosion distribution curves, 313, 315 
progress of erosion, 316 
Austenite in eroded guns 

as solvent for carbon and nitrogen, 
264, 268 

decomposition, 264 
definition, 248 

detection by electron bombardment, 
277 

development, 262-263 
diffraction patterns, 253 
mechanical segregation, 254 
reactivity, 263-264 
removal technique, 257-258 
white layer on gun bores, 247, 264 
Automatic gun mechanism, 538-556 
design for hypervelocity ammuni¬ 
tion, 549-550 

Johnson 20-mm gun, 538-549, 552- 
553 

pre-engraved projectiles, 551-556 
Axial stresses in shells, 523 
A-Z hypervelocity gun, 74-75, 606-608, 
617-618 

Baldwin-Southwark strain gauge, 90 
Ballistic coefficient, formula, 164 
Ballistic pendulum, water, 184-185 
Ballistics of guns, interior, 54-162 
see also Ballistics of hyperveloc¬ 
ity guns, interior 
3-in. gun, 78-94, 135-144 
37-mm gun, 144-148 
band pressure and stress, 152-162 
bore friction, 129-151 
density of loading of powders, 63, 69- 
70 

heating of guns during firing, 98- 
128 

instrumentation for experimental fir¬ 
ings, 76-97 

methods of calculation, 62-66 
powder gas, 21-53 
pre-engraved projectiles, 525-527 
research recommendations, 628-629 
tapered-bore guns, 75 
Ballistics of guns, interior, equations, 
55-62 

basic processes, 55 
burning rates of powders, 62 
equation of energy, 57 
equation of motion, 55-57 
equation of state, 36-38, 57-58 
heat and friction losses, 61-62 
interval after burning, 59-60 
interval of burning, 58-59 


CONFIDENTIAL 


657 


658 


INDEX 


powder constants, 61 
solution, 58-60 

Ballistics of guns, interior, firing test 
measurements, 78-83 
bore friction, 133-135 
ejection of projectile, 82 
firing pin striking primer, 78 
gun recoil, 78, 80 
jump of gun, 82 
muzzle velocity, 79 
photographs of muzzle smoke and 
flash, 82-83 

powder gas pressure, 78, 80-82 
powder gas temperature, 80, 82 
projectile acceleration in gun, 79 
projectile deceleration on range, 79 
projectile travel in gun, 63, 78-79, 
130, 133 

radiation from powder gas, 78 
strain on gun, 82 

Ballistics of guns, interior, instrumen¬ 
tation, 76-97 
accelerometers, 89-90 
Carderock range, 76-77 
comparator for distance measure¬ 
ments on films, 94 
flashmeter, 83, 94 
gauges, 82, 90-92, 95-97 
guns, 77-78 

high-speed cameras, 93-94 
microflash equipment, 83, 94 
microwave interferometer, 92-93, 
133 

optical ejection indicator, 93 
photography, 82-83, 93-94, 274 
projectile gauges, 95-97, 133 
protractor for differentiating curves, 
94-95 

radiation pyrometer, 92, 94 
recoilmeters, 80, 86-87 
recording apparatus, 83-86 
solenoids for projectile velocity, 91 
strain gauges, 82, 90-91 
velocimeters, 80, 86-90 
Ballistics of hypervelocity guns, inte¬ 
rior, 66-75 

design considerations, 70-74 
gun length and muzzle velocity, 74 
limit to muzzle velocity, 67-68, 619 
means of increasing muzzle velocity, 
66 

90-mm gun, 74-75 
optimum conditions, 68-70 
pre-engraved projectiles, 525-527 
research recommendations, 628-629 
sabot-projectiles, 75, 619-624 
subcaliber projectiles, 75, 619-624 
Ballistics of hypervelocity projectiles, 
exterior, 163-179 

comparison with standard projec¬ 
tiles, 163 


deceleration on range, 79, 624 
drag coefficient for a cone moving 
with high velocity, 175-177 
firings of caliber .50 projectiles, 179 
flight characteristics, 163-165 
formula for drag force, 163 
motion of a slightly yawing cone at 
supersonic speeds, 170-175 
photographs of projectiles in flight, 
83 

pre-engraved projectiles, 528-532 
shock waves in air, 83, 170 
stability of subcaliber projectiles, 
165-170 

trajectory determination by tracer 
photography, 177-179 
velocity, 79, 179 

Ballistics of hypervelocity projectiles, 
terminal, 180-192 

armor perforation, 189-192, 626-628 
disruption of a liquid by projectile, 
180-189 
Band, projectile 
see Projectile bands, rotating 
Band pressure of gun tubes, 152-162 
band-to-bore friction, 129-130 
cause of gun erosion, 275 
correlation with bore friction, 153- 
154 

dynamic measurements, 159-162 
evaluation, 135, 157, 522-523 
high-speed strain-recording equip¬ 
ment, 161 

measurements with 37-mm projec¬ 
tiles, 156-159, 161-162 
prediction, 157 
radial stress, 152 
relation to erosion, 159 
static measurements, 156-159 
theory, 152-154 
wear of band, 159 
Barrels (gun) 

aircraft machine gun, 446-453, 458- 
500 

aluminum-clad, 480-481 
cannon, 511-512 
centering cylinder, 266 
cooling methods, 121, 124-128 
Fisa protector, 609-614 
free-run, 471 

heating, 102, 104, 121, 323-325, 466 
heavy machine gun, 223-224, 338, 
414, 453-454 
machine gun, 443-500 
nitralloy, 470 

nitrided, 413, 458-472, 486-491 
special contour, 478 
stellite-lined, 155, 413-414, 443-457, 
473-484 

thermal stresses, 154-156 
velocity measurements, 463 


X-ray examination of erosion, 251 
Barrels, caliber .30; 454-456 
chromium-plated, 471-472 
design and manufacture, 454-455 
erosion-testing gun, 337 
M1919A6 barrel, 455-456 
nitrided and chromium-plated, 471- 
472 

stellite-lined, 454-456, 471-472 
test results, 455-456 
Barrels, caliber .50 
see also Erosion-testing gun 
air-cooling devices, 127 
aircraft, 446-453, 458-500 
band pressure measurements, 162 
chromium-plated, 216, 413-414, 458- 
484 

erosion tests, 223-224, 454 
Fisa protector, 611 
heating rate, 116-117 
heavy barrels, 338, 453-454, 471 
need for improvement, 446-447 
nitrided, 458-472 

Ordnance department tests, 447-451 
pilot plant production, 451-453, 485- 
500 

pre-engraved projectiles, 592,598-599 
rotating projectile bands, 533-534 
stellite-lined, 446-454, 473-484 
thermal stress, 155 
velocity loss, 449 
Barrels, caliber .60 
chromium-plated, 472 
Fisa protector, 609-610, 614 
nitrided, 456-457, 472 
stellite-lined, 456-457 
Barrels, chromium-plated 
see Chromium-plated gun barrels 
Barrels, design, 503-517, 615-617, 630 
erosion-resistant liners, 503-506, 
513-515 

hypervelocity gun, 615-617, 630 
replaceable steel liner, 506-509 
rifling, 515-517 
Barrels, liners for 

see Liners for gun barrels 
Barrels, steel 

see Steel for gun barrels 
Barrels, stresses in 

see Stress in gun barrels 
Battelle laboratory gun, 336 
Beryllium, thermochemical erosion re¬ 
sistance, 354 

Beta-rays, measurement, 242 
Black powder in guns 
effect on cracking of steel, 299 
effect on gun erosion, 297, 299 
Body engraving, 130, 210-212, 516-517 
Bofors 40-mm gun mechanism, modifi¬ 
cations for firing pre-engraved 
projectiles, 553, 555-556 


CONFIDENTIAL 




INDEX 


659 


Bore friction, 129-151 
body engraving, 129 
component forces, 129-130 
definitive equation, 129 
determination from band pressure, 
153-154 

distribution of energy of the powder, 
61-62, 148-151 

effect of powder gas pressure, 130 
effect of temperature, 99-100 
friction coefficient, 157 
friction factor, 599-600 
friction-time curves, 130-131 
heating effect, 100-101,104,116-118, 
135, 318-319 
importance, 129 
pre-engraved projectiles, 594 
reduction, 275, 523 
research recommendations, 629 
Bore friction, determination, 130-135 
average friction, 135 
falling weight tests, 132-133 
from ballistic firings, 133-135 
from band pressure, 153-154 
integration method, 134-135, 147 
measurements in T-47 gun, 144- 
148 

measurements in 3-in. gun, 135- 
144 

microwave interferometer, 133 
projectile acceleration, 133-134 
projectile gauges, 133 
static push tests and strain measure¬ 
ments, 132 

summary of methods, 130-132 
Bore gauge, electrical, 499 
Bore lubricant, 354, 533 
Bore profile, 197 
Bore softening by heat, 261-264 
development of martensite, 263 
heating process, 261-262 
liquefaction of surface, 213-216, 258, 
264, 266, 277 

rapid-fire guns, 123, 262, 475 
single-shot guns, 261-262 
Bore surface, coppering, 217-218 
decoppering methods, 218, 234-235 
definition, 195 
erosion products, 251-254 
friction, 130 

Bore surface, cracking, 276-279, 319 
apparatus for studying, 277 
causes, 274, 276-279 
chromo-plated guns, 213, 410 
entrapped erosion products, 227-228 
patterns, 212-213 
pitting of surface, 277 
plaques, 236 

powder gas erosion, 274, 276-277 
reaction of sulfur gases with gun 
steel, 299-300 


resistance, 335 
shearing of lands, 277 
stress erosion, 277, 279 
Bore surface, erosion 

see Bore surface, cracking; Erosion 
of guns; Erosion products in gun 
bores 

Bore surface, materials 

see Erosion-resistant materials; Steel 
for gun barrels; Stellite, metal¬ 
lurgy and properties 
Bore surface, penetration 

see Carbon penetration into gun steel; 
Nitrogen penetration into gun 
steel 

Bore surface, thermal transformation 
cracking, 277, 279 

critical temperature of steel, 262-263 
distribution of transformed layer, 
319-320 

effect of hydrogen sulfide, 271-272 
martensitic layer, 246 
metallographic examination, 244-246 
rate of formation of transformed 
layer, 320 

relation to erosion, 321, 323-325 
softening of bore, 261-263 
thickness of transformed layer, 320- 
321 

troostite band, 247 
Bore-surface temperature, 105-115 
continued fire, 112-113 
control during firing tests, 224, 286- 
287, 303, 461 

correlation with erosion rate of gun 
steel, 323 

effect of bore-surface material, 114— 
115 

effect of gas leakage, 115, 123 
effect of preheating, 114 
effect on reaction of powder gases 
with gun steel, 250, 269-272, 
286-287, 303, 305-306 
maximum, 118, 121 
melting, 113-114, 257, 333 
Bore-surface temperature, methods of 
determining, 107-112 
comparison of methods, 112 
Fulcher method, 108 
fusion temperature of erosion prod¬ 
ucts, 113-114 

Hicks, Thornhill, and Ware method, 
111 

Hirschfelder, Kershner, and Curtiss 
method, 108-109 
Hobstetter method, 111-112 
Nordheim method, 109-111 
thermocouples, 105-106, 113 
British research 
bore-surface temperature, 111 
heating of gun barrels, 102, 119 


projectile design, 522, 536 
sabot-projectile, 585-587 
Browning machine guns 
see also Barrels, caliber .30; Barrels, 
caliber .50 
cook-off tests, 121 

modifications for firing pre-engraved 
projectiles, 551-552 
performance, 446 
tests with Fisa protector, 613-614 
Bullets 
see Projectiles 

Burning of powder, 21-28, 58-60 
burning constants, 27, 64 
effect of pressure, 35 
effect of radiation, 27-28 
energy released, 38 
equation, 58 

firings in closed vessels, 48-51 
interval after burning, 59-60 
interval of burning, 58-59 
mechanism, 23-24 
pressure-travel curves, 63-65 
rate, 24-28, 62 
stages, 24 

Cadmium-plated projectiles, 354, 535, 
599 

Caliber of guns 

see also Barrels, caliber .30; Barrels, 
caliber .50; Barrels, caliber .60; 
Guns, specific calibers 
effect on heat input to barrels, 99 
effect on muzzle erosion, 206 
effect on muzzle velocity, 622 
effect on origin erosion, 200 
Calorimetric methods of measuring gun 
heating, 102-104 

Cameras, high-speed, 82-83, 93-94, 544 
Cannon, automatic (20-mm) 
see Johnson 20-mm gun 
Cannon tubes 
see Barrels (gun) 

Capacitance-resistance accelerometer, 
90 

Carbides, 253-258 
chromium, 256-257, 438 
formation, 272 
iron, 235-236, 253-257, 268 
molybdenum, 429 
nickel, 257-258 

tungsten, 190-192, 433, 572, 583, 620 
Carbon, radioactive, 33-35, 285 
Carbon equilibrium in powder gas, 33- 
35, 285 

Carbon monoxide, reaction with gun 
steel, 289-296 

ammonia and hydrogen in the charge, 
293 

isolation and identification of iron 
carbonyl, 293-296 


CONFIDENTIAL 





660 


INDEX 


sulfur in the charge, 291-293 
weight losses, 290-291 
X-ray analysis of eroded specimens, 
291 

Carbon penetration into gun steel, 267- 
269, 284-289 

austenite as solvent, 264, 268 
bore-surface temperature, 286-287 
comparison with surface reactions, 
267 

compounds, 284-285 
conditions for penetration, 268 
effect of number of rounds, 288 
effect of propellant, 288-289 
effects of temperature, 286-288 
formation of white layers, 269 
iron-carbon system, 267-268 
method of determining, 285-286 
Carbonaceous materials in eroded gun 
bores, 254 
Carbonyls, 419-424 
chemistry of, 419-420 
chromium, 434-436 
iron, 225, 227, 293-296, 419-420 
molybdenum, 420-425 
nickel, 419-420 
tungsten, 433 
Carburization 

see Carbon penetration into gun steel 
Carderock firing range, 76-77 
Carnegie Institution of Washington 
combination stellite-lined and chro¬ 
mium-plated gun barrels, 473- 
484 

experimental ballistic firings, 76-97, 
159-162, 180-189 
gun barrel liner, 223-224, 338 
nitrided, chromium-plated caliber .50 
aircraft barrel, 458-472 
Casting, methods 
centrifugal, 362 
for stellite, 397, 398, 506 
investment, 397 
vacuum, 359 
Cavitation erosion 
apparatus for studying, 233-234 
tests, 336 

Cementite in eroded guns, 253-257 
disengagement by attacking steel, 
235-236, 255-257 
eutectic temperature, 268 
mechanical segregation, 254 
Centrifugal in-melting, 362-363, 397- 
398, 506 

CGL firing schedule for erosion-testing 
machine gun barrels, 462, 474- 
475 

Chalcocite, detection in eroded guns, 
252 

Chemical alteration of bore surface 
see Erosion products in gun bores 


Chloridizing apparatus for preparing 
molybdenum pentachloride, 421 
Chrome Gage Corporation, pilot plant 
for chromium-plating, 492-499 
barrels handled and processed, 495 
description of project, 492-493 
electrical bore gauge, 499 
electropolishing tank, 493 
plant installation, 493-494 
plating procedure, 495-499 
Chromium, erosion resistance, 331,345, 
356, 358 

Chromium, methods of preparation, 
358-363 

contaminated by nitrogen, 358 
deoxidized with zirconium, 359-360 
electro-deposition, 344,359-360,417- 
418 

impregnations, 363 
inclusions, 358-359 
induction-heated vacuum furnace, 
357 

thermal decomposition of iodide, 360 
vacuum casting, 359-360 
vacuum melting, 358-359 
Chromium, physical properties, 363- 
365 

comparison with gun steel, 363 
ductility, 364-365 

erosion vent-plug tests, 345, 356, 358 
melting point, 331, 358 
Chromium alloys, 360-365 
cobalt, 406 

electroplated coatings, 416-417 
machineability, 365 
nickel-chromium alloys, 351 
physical properties, 363-365 
preparation, 357, 360-363 
research recommendations, 368-369, 
484 

thermochemical erosion resistance, 
345, 367-368, 406-407 
Chromium alloys, gun liners, 364-369 
advantages, 365 
chemical resistance, 368 
disadvantages, 364 
firing tests, 365-367 
mechanical properties, 368 
research recommendations, 368-369 
resistance to powder gas erosion, 
344-345, 351, 367-368 
thermal resistance, 367-368 
Chromium carbide in eroded guns, 256- 
257 

Chromium carbonyl 
formation, 434-436 
Grignard reagent, 434-435 
properties, 436 

Chromium plate, 408-415, 434-439 
see also Electroplated coatings 
annealing, 339, 411 


composition and properties, 411 
duplex coatings, 415, 417 
failure in gun bores, 344, 356, 409- 
410, 601 
hardness, 408 

plating procedures for different base 
materials, 410-411 
research recommendations, 417-418 
surface cracks, 277, 409 
thickness limitations, 408, 600 
types, 410 

vapor-phase deposition, 434-439 
Chromium-plated adaptors for tapered- 
bore guns, 580 

Chromium-plated gun barrels, 411-415, 
458-500 
caliber .30; 471 

caliber .50; 413-414, 458-471 
caliber .60; 472 
disadvantages, 213 
effect of thickness with pre-engraved 
projectiles, 600-601 
electropolishing, 413 
erosion, 213, 216-217, 245, 263-264, 
276-277 

fired with pre-engraved projectiles, 
600-603, 606 

heavy machine gun barrels, 414, 471 
high- and low-contraction chromium, 
412 

methods of plating, 411-413, 459, 
488, 495-499 
naval guns, 356 
90-mm, 606-608, 617-618 
nitrided barrels, 413, 459-460, 465- 
471, 486-491, 495 
pump plating, 414 
specifications, 413, 460, 491 
stellite-lined, 413, 471-484, 491 
temperature-time curves, 461 
thermally-altered surface layer, 245 
Chromium-plated gun barrels, firing 
tests, 460-465 

accuracy measurements, 463-465 
ammunition, 462-463 
firing schedules, 460-462 
Fisa protector, 613 
increase in muzzle velocity and cy¬ 
clic rate, 462 

pre-engraved projectiles, 592,601,603 
targets, 462-464 
temperature-time curves, 461 
velocity measurements, 463 
Chromium-plated gun barrels, perform¬ 
ance, 465-470 
accuracy-life, 465-466 
advantage of choked muzzle, 466 
gauges, 468 
key holing bullets, 466 
projectile engraving, 465 
star gauge curves, 467 


CONFIDENTIAL 




INDEX 


661 


tippers, 468-469 
velocity-life, 466-468 
Chromium-plated gun barrels, pilot 
plants, 485-500 

Chrome Gage Corporation, 492-499 
Doehler-Jarvis Corporation, 485-492 
recommendations, 499-500 
Chromium-plated gun barrels, stellite- 
lined 

see Stellite-lined and chromium-plat¬ 
ed gun barrels 

Chro-mow steel gun barrels, 470, 483 
Chronograph, Aberdeen, 223, 309, 463 
Circumferential strain in gun barrels, 
155-156 

Climax Molybdenum Company 

chromium-base alloys for gun liners, 
356 

molybdenum carbonyl preparation, 
420-425 

thermit melting of molybdenum, 
386-387 

vacuum melting of molybdenum, 
385-386 

Coatings, erosion-resistant 

see Electroplated coatings; Erosion- 
resistant materials; Vapor-phase 
plating 

Cobalt alloys, 351-357 
see also Stellite, metallurgy and 
properties 

hardened iron-nickel-cobalt-chro¬ 
mium alloys, 351-352, 406-407 
molybdenum, 370, 376-377, 383- 
385 

resistance to powder gas erosion, 347- 
348, 351 
steel, 352 

Cobalt gun liners, 347-348, 406 
Cobalt plating, 415-416 
Cobalt-chromium alloy gun liners 
see Stellite gun liners 
Colmonoy gun liner, 350 
Columbium, thermochemical erosion 
resistance, 354 

Comparator for distance measurements 
on films, 94 

Compression of gases, adiabatic 
apparatus, 231-232, 306-307 
maximum temperature and pressure, 
305 

pressure as function of time, 306 
reactions with steel, 305-307 
Contact printing for study of erosion 
products, 240, 252 
Continuous recoilmeter, 87 
Cook-off in guns, 117, 121 
Cooling of gun barrels, 124-128, 481, 
484 

see also Heating of gun barrel during 
firing 


air blasts, 121, 126-127 
convection, 126 
coolants, 127 
devices, 125 

injection sprays, 127-128 
radiation, 126 

rates of cooling, 117, 124, 126 
water jackets, 127 

Copper crusher gauge, use in erosion¬ 
testing gun, 309 

Copper for gun barrels, 353-354 
Copper potassium chloride for re¬ 
moval of erosion products, 235, 
255 

Copper sulfides (gun erosion product), 
252-253 

Coppering of gun bores 

decoppering methods, 218, 234-236 
definition, 195 

entrapped erosion products, 235, 
251-254 

friction, 100, 130, 135 
Corrosion of bore surface 
see Bore surface, thermal transforma¬ 
tion 

Cracking of bore surface 
see Bore surface, cracking 
Crane Company 

gun liners for testing materials, 224, 
338, 443-444 

means of inserting resistant gun 
liners, 443-446, 503-504, 615 
pilot plant for manufacturing stellite- 
lined barrels, 451-453 
replaceable steel liner, 507-512 
Crystal accelerometer, 80, 90 
Crystal gauges, piezoelectric, 96-97 
Cyanides in eroded guns, 249, 253 
Cyanogen in powder gases, 272 
Cyclonite (propellant), 21 
erosiveness, 322, 327 
reaction products, 289 
Cyclops KL gun liners, 352 

Decoppering of gun bores, 218, 234- 
235 

Deformable projectiles 
see also Skirted projectiles 
armor penetration, 191 
De Marre formula for projectile pene¬ 
tration of armor, 190, 587-588, 
626 

Die surfaces, molybdenum-plated, 432 
Diffraction techniques for studying gun 
erosion 

see Electron diffraction, gun erosion 
studies; X-ray diffraction in gun 
erosion studies 
Disks, rupture, 225, 228, 310 
Doehler-Jarvis Corporation, pilot plant 
for chromium-plating, 485-492 


dimensions of nitrided-plated barrels 
with choked muzzles, 489-491 
procedure for nitrided barrels, 486- 
489 

stellite-lined gun barrels, 491-492 
Double-base powders, erosiveness 
see Powders, double-base, erosiveness 
Dowmetal sabot-projectiles, 562 
Drag coefficient, 175-177 
conical-headed projectile, 176 
drag function, 176 
effect of body engraving, 211 
effect of yaw, 175-176 
formula, 163 

function of Mach number, 175 
reduction of drag, 532 
suggestions for further research, 176- 
177 

Draw rifling process for gun tubes, 504 
Duplex coatings, 345, 415, 428 
Dural sabot-projectiles, 562 

E6 thermindex paint, 107 
Eberbach Microvickers hardness tester, 
460 

8-in. howitzer, Mark VII, 124 
Ejection indicator for gun tubes, 82, 93 
Electrical bore gauge, 499 
Electron bombardment, gun erosion 
studies 

apparatus, 232-233 
cracking of bore surface, 277, 335 
formation of austenite, 277 
thermal shock, 277 

Electron diffraction, gun erosion stud¬ 
ies, 237-239 

comparison with X-ray diffraction, 
237 

detection of cyanides, 249, 253 
detection of magnetite, 270 
diffraction patterns, 237 
erosion products from different pow¬ 
ders, 249 

erosion products from sulfur, 298-299 
limitations, 237-238 
microplow, 238-239 
test blocks exposed to powder gases, 
249 

unidentified lines, 239 
vent plugs, 228, 230 
Electronic pulsers, 86 
Electroplated coatings, 408-418 
see also Chromium plate; Vapor- 
phase plating 
alloy deposition, 416-417 
availability of electrodeposited met¬ 
als, 408-409 

characteristics, 408, 418 
cobalt, 415-416 

current distribution during plating, 
498-499 


CONFIDENTIAL 



662 


INDEX 


gun barrels, 411-415, 458-500 
molybdenum, 409, 417 
nickel, 350, 415-416 
pump plating, 414 
requirements, 408 

research recommendations, 417-418 
tantalum, 409 
tungsten, 417 

Electropolishing tank for chromium¬ 
plating, 413, 488, 493 
Elkonite 

projectile nose tips, 583 
resistance to powder gas erosion, 346 
Energy distribution of gun powder, 
148-151 

average temperature, 148 
energy balance during burning, 38, 
149 

energy equation, 148 
heat, 148-149 

kinetic energy, 148, 620-622 
Energy of projectiles 
equation, 57 
in water, 186-187 
kinetic, 148 

loss during flight, 163-164 
loss in gun firing, 61-62, 98-99, 
163 

specific limit energy for armor per¬ 
foration, 189-191 
Engraving of projectiles 
see Pre-engraved projectiles; Projec¬ 
tile engraving 

Epsilon iron nitride in eroded guns, 
255-256 
detection, 249 

m mechanical segregation, 255 
nitrogen content, 268 
stability, 253 
surface layers, 256 
Equations of interior ballistics 
see Ballistics of guns, interior, equa¬ 
tions 

Equivalent stress in gun tubes, 511 
Erosion of guns, 195-327 
asymmetry, 201, 207-208 
correlation with thermal transforma¬ 
tion, 323-325 
definition, 195 

groove erosion, 195, 198, 203 
hypervelocity guns, 11, 13, 16, 615, 
629 

land erosion, 195, 198, 203, 277, 602 
liner joint, 202-203, 255, 274, 403, 
476, 508 

measurement, 197-198 
muzzle erosion, 195, 203-212 
origin erosion, 195, 198-203, 207 
powder gas erosion, 272-275, 283- 
327 

pre-engraved projectiles, 313-317 


relation to muzzle velocity, 10, 200- 
201, 206 

research recommendations, 629 
scoring, 274 

swaging of rifling, 123, 216, 275-276, 
458-459 

Erosion of guns, bore-surface reactions, 
267-272 

see also Bore surface, cracking; Cop¬ 
pering of gun bores 
carbon penetration, 264, 267-269, 
284-289 

equilibrium curves, 271 
hydrogen sulfide, 271-272, 294 
liquefaction of bore surface, 213-216, 
264, 266, 277 

nitrogen compounds, 267-269, 272, 
300-304 

oxidation, 267, 269-270, 283 
pebbling of surface, 215, 278-279, 
319 

softening of the bore, 213-216, 261- 
264, 266, 277 

thermal transformation, 262-264, 
319-325 

water gas components, 269-271 
white layers, 247, 256, 264, 269 
Erosion of guns, causes, 244, 260-279 
bore design, 276 
bore friction, 275 

chemical factors, 260, 267-272, 333 
cracking of bore surface, 276-279 
distortion of rifling, 216-217 
effect of chromium-plate, 263-264, 
410 

gas leakage, 273-274, 592-594 
inclusions, 215 

liquefaction of bore surface, 213-216, 
264, 266, 277 

mechanical, 272-276, 333-334 
mushrooming, 263 
powder composition, 297, 299-300, 
307, 326-327 
principal factors, 260-261 
projectile, 558 
stress, 230, 279 
sulfur, 296-300 

thermal factors, 120-121, 154-156, 
261-267, 277, 513-514 
Erosion of guns, laboratory studies, 
219-243 

see also Erosion products in gun bores 
apparatus for cavitation erosion, 
233-234, 336 

caliber .30 barrel, 286, 294-295, 297, 
302 

caliber .50 erosion-testing gun, 220- 
223, 309-310, 336-337 
caliber .50 machine guns, 223-224 
chemical analysis, 239-240 
collection of particles and gases dis¬ 


charged from gun barrels, 230- 
231 

determination of the melting tem¬ 
peratures of erosion products, 
240 

electron bombardment apparatus, 
232-233, 277, 335 

electron diffraction, 228-230, 237- 
239, 249, 270, 298-299 
firing tests, 220-231 
gauges, 197-198, 222-223, 225 
isotopic tracers, 240-243,, 285, 297 
metallographic examination of eroded 
surfaces, 236-237, 244-246, 319- 
321 

vent plugs, 224-232, 310-311, 322- 
323, 334-335 

visual examination of eroded sur¬ 
faces, 236-237 

X-ray diffraction, 237-239, 248-251, 
291, 301 

Erosion of guns, mitigation 
see also Liners for gun barrels 
electroplated coatings, 408-418, 460, 
488 

Fisa protector, 609-614 
modification of powders, 283, 300, 
307, 325-327 

pre-engraved projectiles, 591-608 
requirements, 408 
test of ferrosilicon, 327 
thermodynamic considerations, 326- 
327 

Erosion products in gun bores, 244- 
259, 321-322 

collection of gases and particles, 230, 
249, 301, 335 

comparison of single- and double¬ 
base powders, 248-251 
entrapped in the coppering, 251-254 
fusion temperature, 113-114, 240, 
258 

melting, 257 

vinylite coatings for examination, 
236 

Erosion products in gun bores, removal 
techniques, 234-236, 254-258 
by attacking the steel, 235-236 
chemical removal, 255-258 
copper potassium chloride solution, 
235, 255 

coppering and entrapped products, 
234-235, 251-254 
degreasing of test specimens, 234 
fractionation of erosion products, 236 
mechanical removal, 234, 254-255 

Erosion products in gun bores, specific 
substances 
alabandite, 236 

austenite, 247-249, 253-254, 257- 
258, 262-264 


CONFIDENTIAL 




INDEX 


663 


cementite, 235-236, 253-257, 268 
chromium carbide, 256-257 
cyanides, 249, 253 

epsilon iron nitride, 249, 253, 255- 
256, 268 
ferrite, 248-249 
galena, 252 

iron carbides, 253-257, 268 
iron carbonyl, 225, 227, 293-296, 
419-420 

iron nitrides, 253, 255-256, 268 
iron oxides, 273 
lead, 252 

magnetite, 254, 270, 306 
martensite, 263 
nickel carbide, 257-258 
nitrides, 255-256, 268 
potassium sulfate, 252 
pyrrohotite, 254 
sphalerite, 253 
wustite, 249, 253 

Erosion products in gun bores, surface 
layers, 244-248, 255-258 
austenite-rich layers, 257, 262-264 
cementite-rich layers, 256 
characteristics, 244 
chromium-plated guns, 245 
constituents, 255 
effect of single-base powders, 247 
inner and outer white layer, 244, 247- 
248, 258, 264 

liquefaction of bore surface, 213-216, 
258, 264, 266, 277 
martensite layer, 246, 263 
techniques of studying, 234-240, 
244 

thermally-altered layer, 245-246, 
258-259, 261-264 
troostite band, 247 
Erosion studies in vents, 220-232 
see also Erosion vent-plug tests 
apparatus, 228, 230 
D-shaped cross section, 227-230 
method, 230, 334-335 
results of studies, 249, 272-275, 277- 
279, 299-300 

rifled barrel as vent, 225-227, 294 
stressed vents, 230, 272-275 
studies of cracking, 227-228 
use of oxygen and carbon monoxide 
mixtures, 227 

Erosion vent-plug tests, 224-230 
apparatus, 224-225 
carbon monoxide reactions, 290-293 
chromium, 334, 345, 356, 358 
detection of iron carbonyl, 225, 227, 
294 

powders, 310-311, 322-323 
resistance of metals and alloys, 334- 
335 

rupture disks, 225, 228, 310 


subjected to adiabatically com¬ 
pressed gases, 231-232, 305-307 
X-ray diffraction examination, 228, 
230 

Erosion-resistant materials, 331-439 
applications, 343-355, 440-506, 513- 
515 

chemical resistance, 333 
chromium and alloys, 263-264, 344- 
345, 351-352, 356-369, 416-417 
cobalt alloys, 391-407 
desirable properties, 333-334 
electrodeposited coatings, 408-418 
evaluation, 343-355 
fluorocarbon coating, 354 
machineability, 343, 371, 396 
melting points, 333 
molybdenum, 334, 370-390 
nickel alloys, 334, 349-351, 406-407 
parco-lubrite coating, 354 
research program, 331-332 
research recommendations, 629 
steels, 337 

stellite, 334, 391-407 
tantalum, 334, 347 
thermal properties, 333, 338-340, 
513-514 
tungsten, 334 
types, 331 

vapor-phase plating, 419-439 
Erosion-resistant materials, mechanical 
properties, 340-343 
adherence of coatings, 343 
flaw detection, 343 
hardness, hot-hardness, and wear re¬ 
sistance, 342 
machineability, 343 
requirements, 333-334 
tensile strength, ductility, and modu¬ 
lus of elasticity, 340 
Erosion-resistant materials, test meth¬ 
ods, 334-343 

Battelle laboratory gun, 336 
bend and ring-compression tests, 343 
caliber .30 machine gun liners, 337 
caliber .50 erosion-testing gun, 336- 
337 

caliber .50 machine gun liners, 338 
cavitation erosion tests, 336 
erosion vent plugs, 334-335 
firing of metal particles into an evac¬ 
uated tube, 335-336 
hydrostatic tube testing, 342-343 
notch impact tests and high velocity 
impact tests, 340-342 
resistance to surface cracking, 335 
Erosion-testing gun, caliber .30; 337 
Erosion-testing gun, caliber .50; 219- 
223 

applications, 219, 222-223, 336-337 
design, 220, 309, 336-337 


limitations, 223 
molybdenum liners, 377-385 
tests on Fisa protector, 612-614 
tests using pre-engraved projectiles, 
313-317, 592-600 
use of gauges, 222, 309 
X-ray examination of barrels, 251, 
310, 321 

Erosion-testing gun, propellant tests, 
313-322 

artillery-type bullets, 313, 317 
chemical products on the bore sur¬ 
face, 321-322 
conditions of test, 309 
distribution of erosion, 313-315 
erosion rate, 317-318 
heat input to the barrel, 318-319 
measurements, 309-310 
metallographic changes accompany¬ 
ing firing, 319-321 
pre-engraved bullets, 313-317 
progress of erosion, 313, 316-317 
X-ray diffraction examination of bar¬ 
rels, 251, 310, 321 
Erosive action of powders 
see Powders, erosiveness 
Exelloy gun steel, 484 
Explosion vessel for vent-plug tests, 
224-225 

Explosives for propellants, 21 
Exterior ballistics of projectiles 

see Ballistics of hypervelocity pro¬ 
jectiles, exterior 

Falling weight apparatus for determin¬ 
ing bore friction, 100, 132-133 
Ferrite (erosion product), 248 
Ferrosilicon 

reduction of gun erosion, 327 
resistance to chemical attack, 336 
.50 caliber guns 
see Barrels, caliber .50 
57-mm gun, erosion products, 255 
57/40-mm tapered-bore guns 
accuracy tests, 584-585 
design, 570-572 

design of skirted projectile, 572 
muzzle adaptor, 571 
wear after firing, 580 
57/40-mm tapered-bore guns, armor 
penetration tests, 585-590 
ballistic limits of projectiles, 585-587 
DeMarre formula, 587-588 
design changes to improve penetra¬ 
tion, 585 

perforation range comparisons, 588- 
590 

Firing schedules for testing gun barrels, 
309, 460-462, 474-475 
Fisa protector for gun barrels, 609-614 
caliber .50; 612-613 


CONFIDENTIAL 



664 


INDEX 


caliber .60; 609, 614 
design, 14, 610-612 
effect on velocity-life, 609 
extent of engraving, 611 
manufacture, 612 
tests, 612-614 
37-mm, 614 
5-in. gun, 253 

Flame temperature of powders 

see Adiabatic flame temperature of 
powders 

Flashmeter, measurement of muzzle 
flash intensity, 83, 94 
Flight characteristics of hypervelocity 
projectiles, 163-165 
ballistic coefficient, 164 
range and time of flight at high 
speeds, 164-165, 624-626 
subcaliber projectiles, 165, 619-626 
velocity and energy loss, 163-164, 
619-623 

Fluorocarbon for coating gun bores, 
354 

FNH-M1 gun powder 
effect on gun erosion, 250 
erosion products, 270 
temperature of powder gas, 122 
FNH-M2 gun powder, erosiveness, 250, 
270, 289 

Forcing cone, 158, 198, 595 
Formulas 

ballistic coefficient of projectile, 164 
burning of gun powder, 58 
De Marre formula for projectile 
penetration of armor, 190, 587- 
588, 626 

drag coefficient of projectile, 163 
energy distribution of gun powder, 
148 

energy of projectile, 57 
Lame formulas for gun tube stresses, 
152, 511 

moment coefficient of projectile, 
167 

motion of projectile, 55-57 
muzzle velocity, 63, 67-68, 620-624 
projectile design, 521-522 
stability of subcaliber projectiles, 168 
stresses in gun barrels, 511 
4.7-in. gun 

origin erosion, 199 
star-gauge curves, 196 
14-in. gun, 203, 212, 253 
40-mm guns, 118, 121, 553-556 
Franklin Institute 

caliber .50 erosion-testing gun, 220- 
223, 336-337, 377-385 
Fisa protector, 609-614 
"Free-run” machine gun barrels, 471 
Friction in guns 
see Bore friction 


G8-type projectile, 163-164 
Galena, detection in eroded guns, 252 
Gamma iron, 248 
see also Austenite in eroded guns 
Gamma-prime iron nitride in eroded 
guns, 255-256, 268 
Gas compression, adiabatic 
apparatus, 231-232, 306-307 
maximum temperature and pressure, 
305 

pressure as function of time, 306 
studies on powder gas reaction with 
steel, 305-307 

Gas leakage in guns, 592-596 
bore surface temperature, 115, 123 
causes, 115 

erosive effect, 273-274, 562, 592-594 
pre-engraved projectiles, 596 
pressure, 202 
sabot-projectile, 562 
swaging of rifling, 123 
velocity performance, 202, 595-596 
Gas pressure shock waves, suppression 
in guns, 576-578 

Gas turbulence in guns, 53, 266, 273, 
613 

Gas-operated machine gun 
see Johnson 20-mm gun 
Gauges 

electrical bore gauge, 499 
erosion gauge, 196-198, 222-223, 225 
liner rejection gauge, 468 
plug, 198, 222, 309-310, 469-470 
Sheffield Precisionaire gauge, 487 
star gauge, 196-198, 222, 467 
thread-collocating gauge for machin¬ 
ing gun tubes, 578 
use in production of chromium-plated 
gun barrels, 486-488, 499 
Gauges for ballistic measurements, 90- 
92, 95-97 

bore friction measurements, 133 
copper crusher, 225, 309 
piezoelectric, 96-97 
pressure, 91-92, 309 
projectile gauges, 95-97, 133 
resistance, 91-92 
strain, 82, 90-91 

Geiger-Mueller counters, 240-242, 296- 
297 

detection of radioactive carbon, 285 
detection of sulfur in eroded guns, 
240-241, 296 
gun erosion studies, 242 
use of rifle barrel, 297 
Geophysical Laboratory, C.I.W. 
combination stellite-lined and chro¬ 
mium-plated gun barrels, 473- 
484 

experimental ballistic firings, 76-97, 
159-162, 180-189 


gun barrel liner, 223-224, 338 
nitrided, chromium-plated caliber .50 
aircraft barrel, 458-472 
GL-135 firing schedule for testing heavy 
barrels, 461 

Gold-palladium alloy, thermochemical 
erosion resistance, 354-355 
Greenough stereoscopic microscope, 236 
Grignard reagent for producing chro¬ 
mium carbonyl, 434-436 
Groove erosion in gun bores 
at muzzle, 203 
at origin of rifling, 198 
definition, 195 
Gun ballistics 

see Ballistics of guns 
Gun barrels 

see Barrels (gun) 

Gun bores 

see Bore friction; Bore softening by 
heat; Bore surface; Heat input 
to bore surface 
Gun chamber 
fluting, 457 

integral with liner, 457 
Gun erosion 

see Erosion of guns 
Gun firing, thermal effects, 98-128 
see also Bore-surface temperature; 
Cooling of gun barrels; Heat in¬ 
put to bore surface; Heating of 
gun barrel during firing; Ther¬ 
mal stresses in gun barrels 
accuracy drop, 121-122 
ballistics, 121-123 
bore friction, 99-100 
burning rate of powders, 27 
erosion, 120-121, 308 
expansion of bore, 123, 466, 476-477, 
480 

pressure and velocity changes, 122- 
123 

softening of bore surface, 213-218, 
234-235, 251-254, 261-264, 276- 
279 

weakening of barrel wall, 123-124, 
476-477, 480 
Gun friction 
see Bore friction 
Gun liners 

see Liners for gun barrels 
Gun mechanism, automatic, 538-556 
design for hypervelocity ammunition, 
549-550 

Johnson 20-mm gun, 538-549 
pre-engraved projectiles, 551-556 
Gun powders 
see Powders 

Gun recoil measurements 
accelerometers, 80, 89-90 
ballistic firing tests, 78, 80 


CONFIDENTIAL 




INDEX 


665 


displacement of gun, 80 
pressure in recoil cylinders, 82 
recoilless gun blocks, 431-432 
recoilmeters, 80, 86-87 
starting time, 78 
velocimeters, 80, 88-89 
Gun steel 

see Steel for gun barrels 
Gun tubes 

see Barrels (gun) 

Guns, coppering 

decoppering methods, 218, 234-236 
definition, 195 

entrapped erosion products, 235, 
251-254 

friction, 100, 130, 135 
Guns, general types 
air-cooled, 121, 127 
antiaircraft, 124 

for firing pre-engraved projectiles, 
551-553, 555-556, 591-608 
howitzers, 124, 563-564, 566 
tapered-bore, 15, 75, 569-590 
water-cooled, 121, 127 
Guns, pressures in 
see Pressure in guns 
Guns, rifling design 
see Rifling of guns 
Guns, specific calibers 

see also Erosion-testing gun 
.30 caliber, 337, 454-456, 471-472 
.50 caliber; see Barrels, caliber .50 
.60 caliber, 456-457, 472, 609-610, 
614 

3-in.; see 3-in. gun 

3.7- in., 119, 121, 249 

4.7- in., 196, 199 
5-in., 253 
8-in., 124 

14-in., 203, 212, 253 
20-mm Hispano-Suiza, 118, 121, 

133 

20-mm Johnson automatic, 123, 538- 
549 

37-mm; see 37-mm guns 
40-mm, 118, 121 
57-mm, 255 

57/40-mm tapered-bore, 570-572, 
580, 584-590 

75- mm, 119, 156-159, 535-536, 563- 

564 

76- mm, 122, 214, 564-565 
90-mm, 74-75, 507-510, 564-567, 

606-608, 617-618 
105-mm, 118-119, 124, 566 
155-mm, 128, 201, 210 
Bofors, 40-mm, 553, 555-556 
Browning machine gun, 121, 212, 
551-552, 613-614 
Hispano-Suiza, 118, 121, 133 
Johnson Ml945; 455 


Hastelloys 

gun liners, 350 

thermochemical erosion resistance, 
407 

Haynes Stellite Company, investment- 
cast stellite liners, 447 
Heat input to bore surface 

see also Bore-surface temperature; 
Heating of gun barrel during 
firing 

bore friction as source of heat, 61, 
100-101, 104, 116-118, 318- 

319 

correlation with erosiveness of pow¬ 
der, 324 

correlation with flame temperature of 
powder, 318-319, 324 
powder gases as source of heat, 62, 
99, 101-102, 149 
radiation as source of heat, 99 
37-mm gun, T47; 101 
3-in. gun, Mark 3; 101 
variation with caliber, 109 
Heat input to bore surface, methods of 
measurement, 102-107 
calorimetric methods, 102-104 
embedded thermocouple, 105-106, 
113 

heating by firing at steady rate, 102- 
103 

strain measurements, 82, 106 
surface-temperature measurements, 
104-105 

thermal analyzer, 107 
Heat losses in guns, 61-62 
Heat of explosion of powder gas 
see Powder gas, thermodynamics 
Heat transfer coefficient, 100 
Heating of gun barrel during firing 
see also Bore-surface temperature; 
Cooling of gun barrels; Heat in¬ 
put to bore surface; Thermal 
stresses in gun barrels 
ballistic equations, 61-62 
effects, 120-124, 216 
rate of heating, 115-120 
temperature-time curves for ma¬ 
chine-gun barrels, 116-117, 461 
Heavy machine gun barrel 
caliber .50; 338, 453-454, 471 
chromium-plated, 414 
liners for, 223-224 
HI A grade 160 gun steel, 483 
High-speed cameras, 82-83, 93-94 
High-speed machine gun, 478 
Hispano-Suiza gun 

bore-friction measurements, 133 
improvement, 538 
Mark V (20-mm), 118, 121 
sabot-projectiles, 562-563 
Hoop stress in shells, 152, 523 


Howitzers 
8-in. Mark VII, 124 
75-mm, 563-564 
105-mm, 566 
Hydraulic obturator, 96 
Hydrocyanic acid in powder gases, 272 
Hydrogen in eroded guns 

reaction with carbon monoxide, 293 
reaction with steel, 274-275 
Hydrogen sulfide, 292-294 
effect on bore surface, 271-272 
effect on erosiveness of carbon mix¬ 
tures, 292-293 
effect on gun erosion, 294 
molybdenum sulfide plates, 427 
reaction with steel, 296 
Hydrostatic tube-testing device, 342- 
343 

Hypervelocity guns 
automatic mechanism, 549-556 
A-Z gun, 74-75, 606-608, 617-618 
definition of hypervelocity, 8 
erosion, 11, 615 

erosion-resistant liners and coatings, 
14 

Fisa protector, 14, 609-614 
gun tube design, 503-517, 630 
interior ballistics, 66-75, 628-629 
molybdenum liners, 382-385, 6lb- 
617 

replaceable steel liners, 14, 507 
research recommendations, 628-630 
rifling design, 515-516 
stellite liners, 405, 615-616 
summary of volume, 1-4 
tapered-bore, 15, 75, 569-590 
Hypervelocity guns, ballistic design 
considerations, 70-74 
choice of powder, 73-74 
limiting density of loading, 70-71 
limiting pressure, 71-72 
90-mm gun, 74 
preferred volume ratio, 72 
Hypervelocity projectiles, 9-15, 163- 
170, 557-608 
see also Projectiles 
advantages, 9-10, 618-628 
comparison with standard, 163 
drag coefficient, 175-177 
flight characteristics, 163-165 
limitations, 10-12 

pre-engraved, 313-317,523-533,551- 
556, 591-608, 630 
range, 9, 164-165, 202 
sabot, 169-170, 557-568, 585-587, 
630 

skirted, 569-590, 620, 630 
subcaliber, 165-170, 618-628 
summary of volume, 1-4 
tungsten carbide-cored, see Tungsten 
carbide-cored projectiles 


CONFIDENTIAL 



666 


INDEX 


Hypervelocity projectiles, armor per¬ 
foration 

see Armor-piercing projectiles 
Hypervelocity projectiles, ballistics 
see Ballistics of hypervelocity pro¬ 
jectiles 

Hypervelocity projectiles, design, 518- 
537 

British, 522, 536 
core, 572, 583 

low-stress rotating bands, 533-536 
materials used, 192 
methods for detecting defects in shells, 
536-537 

pre-engraved, 523-525 
sabot, 559-562 
stresses in shells, 518-524 
Hypervelocity projectiles, effect of yaw, 
170-175 

drag coefficient, 175-176 

normal force and center of pressure, 

172- 174 

normal head force coefficient and yaw 
ratio, 174-175 

ratio of shock yaw to projectile yaw, 

173- 174 

recommendations for future research, 

177 

theory, 170-172 

Hypervelocity projectiles, liquid pen¬ 
etration, 180-189 
effect of velocity, 180 
firings into cans of liquid, 180-181 
motion pictures of jets during disrup¬ 
tion, 181-183 

pressures in water, 184-187 
research recommendations, 187 
shock waves, 181-183, 187-188 
transfer of momenta, 188-189 

Injection sprays for cooling gun barrels, 
127-128 

Interferometer, microwave, 92-93, 133 
Interior ballistics 

see Ballistics of guns, interior; Bal¬ 
listics of hypervelocity guns, 
interior 

Iridium, thermochemical erosion re¬ 
sistance, 355 
Iron alloys 
see also Steel 

thermochemical erosion resistance, 
331, 406-407 

Iron carbide in eroded guns, 253-257 
disengagement by attacking steel, 
235-236, 255-257 
eutectic temperature, 268 
mechanical segregation, 254 
Iron carbonyl from eroded guns, de¬ 
tection, 293-296 
chemistry of, 419-420 


effect of hydrogen sulfide, 294 
experimental method, 293 
formation, 225, 227 
gas mixture in caliber .30 barrels, 294 
location of erosion products, 293-294 
solid iron-bearing deposits, 295-296 
vent plugs, 225, 227, 294 
Iron nitrides in eroded guns 
epsilon iron nitride, 249, 253, 256, 268 
formation, 268 

gamma-prime iron nitride, 255-256, 
268 

mechanical segregation, 255 
resistance, 253 

Iron oxides in eroded guns, 273 
Iron sulfide in eroded guns, 254 
Iron-carbon system, 267-268 
Iron-nitrogen system, 268 
Isothermal transformation diagram for 
steel, 275 

Isotopic tracers, gun erosion studies, 
240-243, 285, 296 

demonstration of carbon equilibrium, 
240-241 

Geiger-Mueller counters, 242 
sampling, 241-242 
temperature of specimen, 242-243 

Jet propulsion motor, radiation and 
temperature of gases, 51-52 
Johnson 20-mm gun, 538-549 

automatic chambering mechanism, 
540-541, 552-553 

case ejection and cocking failure, 544 
description of models, 539-542 
firing rate, 545-549 
kinematic analyses, 544-545 
Navy specifications, 538-539 
parts breakage, 543-544 
“pulled” projectiles, 543 
rebound of impacting parts, 543 
recommendations, 548-549 
Jump ratio of projectiles, 526 

Keyholing projectiles, 450, 466 
Kinetic energy of projectiles, 148 

Lame formulas for gun tube stresses, 
152, 511 

Land erosion in gun bores 
at muzzle, 203 
at origin of rifling, 198 
definition, 195 
effect on gun life, 602 
shearing of lands, 277 
Lead in eroded guns, detection, 252 
Leeds and Northrup Company, heat 
measurements for guns, 104 
Linear velocimeter, 89 
Liners for gun barrels, 440-457 
Cyclops KL steel liners, 352 


erosion, 202-203, 255, 274, 403, 509 
investment-cast stellite, 397, 447, 615 
90-mm gun, 507-510 
protrusion of, 124 
replaceable steel liners, 14, 507-510 
Liners for gun barrel's, design, 503-509 
bonding of liner to barrel, 506 
Crane Company liner, 224, 338, 443- 
444 

effect of differences in thermal prop¬ 
erties, 513-514 

expansion of the liner by draw rifling, 
504 

Geophysical Laboratory liner, 223- 
224, 338, 473-484 

mechanical retention of the liner, 
504-506 

methods of inserting, 366-367, 443, 
503-509 

research recommendations, 368-369, 
484, 616-617, 629 

shrink fit of tube and liner, 503-504, 

512 

stresses in, 512-515 

support of brittle liner materials, 

513 

Liners for gun barrels, materials tested, 
343-355 

characteristics, 331, 351 
chromium and alloys, 344-345, 351- 
352, 364-369 

cobalt and alloys, 347-348, 351-352, 
391-407 

copper and alloys, 353-354 
evaluation, 351-352 
hardened alloys, 350-351 
molybdenum, 345-346, 370-390, 616- 
617 

monel-metal, 350 

nickel and alloys, 349-352, 406-407 
Refractaloy No. 70; 351-352 
requirements, 344 

research recommendations, 617, 629 
steel, 14, 351-353, 506-509, 513 
stellite, 351, 391-407, 443-457, 473- 
484 

tantalum, 347 

TEW alloy, 335, 352, 407 

tungsten, 346-347 

Link-Belt Company, nitriding of steel, 
459-460 

Longitudinal stresses in shells, 153, 
524-525 

Lubrication of gun bore, 354, 533 

Mach number, relation to drag coeffi¬ 
cient, 175 
Machine guns 

see Browning machine guns; Guns 
Magnetite in eroded guns 

detection by electron diffraction, 270 


CONFIDENTIAL 





INDEX 


667 


formation, 306 
mechanical segregation, 254 
Magnetron anodes, molybdenum-plat¬ 
ed, 432 

Martensite in eroded guns, 263 
Martensitic layer, 246 
Massachusetts Institute of Technology, 
strain-recording equipment, 161 
Metals, plating 

see Electroplated coatings; Vapor- 
phase plating 

Metals, thermochemical erosion resist¬ 
ance 

see Thermochemical erosion resist¬ 
ance of metals and alloys 
Methane in quenched powder gas, 32 
Microflash equipment, 83, 94 
Microplow for electron diffraction 
measurements, 237-239 
Microscope, stereoscopic, 236 
Microwave interferometer, 92-93, 133 
Mises-Hencky theory, 152, 510, 522 
Molybdenum, 370-390 
bend strength and yield, 374 
carbides, 429 
chloride, 420-421 
coatings, 388-390, 409, 420-433 
commercial, 370-373 
effect of temperature on hardness, 
376 

hardening by alloying, 375-376 
Molybdenum, methods of preparation, 
385-389 

commercial process, 370-371 
hydrostatic compression test, 375 
inductive sintering in an atmosphere 
of hydrogen, 388-389 
melting in thorium oxide refractories, 
387 

thermit meltings, 387-388 
vacuum melting in an electric arc, 
385-386 

vacuum melting in an induction fur¬ 
nace, 386-387 
working schedule, 373-375 
Molybdenum, vapor-phase plating, 
420-433 
apparatus, 426 
applications, 430-432 
general procedure, 424-425 
plate-to-metal bond, 428-430 
properties of plates, 425-428 
recommendations, 432-433 
sulfide plates, 427, 432 
Molybdenum alloys, 374-377, 383-387 
cobalt, 370, 376-377, 383-385 
electroplated coatings, 417 
hardness, 376 

melting alloys in aluminum oxide 
crucible, 386-387 
nickel, 376-377 


tungsten, 376, 434 
vacuum-melting furnace, 385-386 
working schedule, 374 
Molybdenum carbonyl, 420-425 
drying and sublimation, 423-424 
flow-sheet for production, 425 
formation, 420 

preparation of anhydrous molybde¬ 
num chloride, 420-421 
preparation of crude product, 421- 
422 

properties of sublimed product, 423- 
424 

recommendations, 424 
removal of ether, 422-423 
steam distillation, 423 
Molybdenum gun liners 
alloy with cobalt, 383-385 
caliber .50 erosion-testing gun, 377- 
385 

causes of failure in tests, 382-385 
commercial molybdenum, 373 
firing tests, 382-385 
insertion, 514 
requirements, 431 
suitability of alloys, 346 
thermal properties, 513-514 
3-in. gun, 616-617 

Molybdenum gun liners, method of 
fabrication 

brazed-in one-seam tube, 381 
cast-bonded molybdenum-steel strip, 
381-382 

design considerations, 377 
for caliber .50 erosion-testing gun, 
377-385 

seamless tube, 382 
tube made from flat disks, 381 
wire-wrapped tube, 381 
Molybdenum gun liners, stave-type, 
377-381 

construction, 378-379 
deficiencies in design, 380-381 
milling operation, 383 
precision twisting, 379, 382, 384 
seams, 378 

stresses, 377-378, 514-515 
troughing operation, 380 
Molybdenum pentachloride, apparatus 
for preparing, 421 

Molybdenum steel for machine gun 
barrels, 353, 381-382, 470, 482- 
483 

Molybdenum-tungsten alloy plates, 434 
Moment coefficient, formula, 167 
Monel-metal gun liners, 350 
Motion of projectile, equation, 55-57 
Mushrooming (pocketlike erosion), 263 
Muzzle adaptors for 57/40-mm tapered- 
bore guns, 571, 575-580 
chromium-plated, 580 


design, 575-576 
tests, 579 
ven^s, 578 

wear after firing, 579 
Muzzle blast 
effect of gun length, 72 
photography of flash, 82-83 
Muzzle erosion, 203-212 
asymmetry of erosion, 207-208 
correlation with body engraving, 
210-212 

correlation with origin erosion, 207 
definition, 195 
dependence on caliber, 206 
dependence on muzzle velocity, 206 
development with firing, 204-206 
effect on gun performance, 208-211 
effect on projectile yaw, 211 
land and groove erosion, 203 
variation along the bore, 203-204 
Muzzle flash 
effect of gun length, 72 
photography, 82-83, 93-94 
Muzzle velocity 

dependence on gun caliber, 622 
effect of gun length, 74 
effect on armor perforation, 192 
formula, 63 

means of increasing, 66, 549-550 
measurement, 79 
relation to erosion, 200-201, 206 
subcaliber projectiles, 622 
upper limit, 67-68, 619 

National Bureau of Standards 
experimental ballistic firings, 76-97, 
159-162 

nitrided, chromium-plated caliber .50 
aircraft barrel, 458-472 

Nickel 

alloys, 350-352, 376-377, 406-407 
carbide, 257-258 
carbonyl, 419-420 
gun liners, 349-352 
plating, 350, 415-416 
90-mm guns 

chromium-plated, 606-608 
cooling, 128 
design, 74-75, 606-608 
erosion, 249, 509-510 
molybdenum liner, 616-617 
replaceable steel liner, 507-509 
sabot-projectiles, 564-567 
Nitralloy for machine gun barrels, 470 
Nitric oxide in quenched powder gas, 
32 

Nitrided and chromium-plated gun bar¬ 
rels, 458-472 

accuracy and velocity-life, 458 
caliber .60; 456-457, 472 
hardness profile, 460 


CONFIDENTIAL 



668 


INDEX 


method of nitriding, 413, 459-460 
nitralloy, 470 
specifications, 459-460 
stellite-lined and chromium-plated, 
401-402, 491-492 
temperature-time curves, 461 
velocity-life, 458, 466-468 
Nitrided and chromium-plated gun bar¬ 
rels, firing tests, 460-465 
accuracy measurements, 463-465 
ammunition, 462-463 
change in muzzle velocity versus cy¬ 
clic rate, 462 
firing schedules, 460-462 
targets, 462-464 
temperature-time curves, 461 
velocity measurements, 463 
Nitrided and chromium-plated gun bar¬ 
rels, performance, 465-470 
accuracy-life, 465-466 
advantage of choked muzzle, 466 
appearance of “tippers”, 468-469 
gauge for indicating rejection point, 
469-470 

keyholing bullets, 466 
projectile engraving, 465 
star gauge curves, 467 
velocity-life, 466-468 
Nitrided and chromium-plated gun bar¬ 
rels, plating process, 413, 486- 
491 

anodic etch prior to plating, 488 
caliber .50 heavy barrels, 471 
cleaning operations, 487 
electroplating, 488 
electropolishing, 488 
gauging operations, 487-488 
plating procedure, 486-487 
with choked muzzles, 489-491 
Nitrided gun liners, 353, 401-402 
Nitrides in eroded guns 
epsilon iron nitride, 249, 253, 255- 
256, 268 

formation, 268, 272, 301-303 
gamma-prime iron nitride, 255-256, 
268 

mechanical segregation, 255 
Nitroaustenite, 268 
see also Austenite in eroded guns 
Nitrocellulose, thermal decomposition, 
23-24 

Nitrogen penetration into gun steel, 
267-269, 300-304 
austenite as solvent, 268 
bore-surface temperatures, 303 
comparison with surface reactions, 
267 

conditions for penetration, 268 
cracking, 300 
effect of ammonia, 301 
formation of nitrides, 272, 301-303 


formation of white layers, 269 
iron-nitrogen system, 268 
method of determining, 303 
nitrogen content of reaction products, 
304 

Nitroglycerin 

in gun powders, 327 
thermal decomposition, 23-24 

Obturation 

see Gas leakage in guns 
Obturator, hydraulic, 96 
Occlusion of gases, 274-275 
105-mm guns 
antiaircraft, 124 
heating rates, 118-119 
105-mm howitzer, 566 
155-mm guns, 128, 201, 210 
Optical ejection indicator, 93 
Origin erosion in gun bores, 198-203 
asymmetry of erosion, 201 
correlation with muzzle erosion, 207 
definition, 195 
dependence on caliber, 200 
dependence on muzzle velocity, 200- 
201 

development with firing, 199-200 
effect on gun performance, 201-202 
rate, 201 

variation along the bore, 199 
Oscillograph for ballistic measurements, 
83-86 

Oxygen, reaction with steel, 269-271 
Oxygen compressor cylinders, molyb¬ 
denum-plated, 432 

Paints, temperature-indicating, 106- 
107 

Palladium-gold alloy, thermochemical 
erosion resistance, 354-355 
Parco-Lubrite for coating gun bores, 
354 

Parco-lubrized projectiles, 275, 599 
Parco-lubrizing process, 533 
PE projectiles 
see Pre-engraved projectiles 
Pebbling of gun bore surface, 215, 278- 
279, 319 

Peerless A gun steel, 483 
Pendulum, water ballistic, 184-185 
Phenyl magnesium bromide, use in pro¬ 
ducing chromium carbonyl, 434- 
436 

Photoelectric pyrometer 
apparatus, 46-47 
design principles, 46 
3-in. gun measurements, 92, 94 
Photography in ballistics 
gas leakage, 274 

high-speed cameras, 82-83, 93-94 
muzzle flash, 82-83, 93-94 


muzzle smoke, 82-83 
projectile disruption of water, 181— 
183 

projectile in flight, 83, 177-179 
Piezoelectric crystal gauges, 96-97 
Pilot plants for chromium-plating of 
gun barrels, 485-500 
Chrome Gage Corporation, 492-499 
Doehler-Jarvis Corporation, 485- 
492 

recommendations, 499-500 
Pilot plants for stellite-lined gun bar¬ 
rels, 451-453 

Plaques for studying cracks in gun 
bores, 236 

Plastics for sabot-projectiles, 562 
Plating 

see Chromium plate; Electroplated 
coatings; Vapor-phase plating 
Plug gauges 

erosion measurements, 198, 222, 309- 
310 

indicating rejection point of gun bar¬ 
rels, 469-470 

Potassium sulfate in eroded guns, 
252 

Potomac die-steel for machine gun bar¬ 
rels, 470, 483-484 
Powder gas, 21-53 
see also Gas 
covolume, 36, 57 

heat transfer to gun, 62, 99-102, 149 
molecular constituents, 37 
turbulence, 53, 266 

Powder gas, chemical equilibrium, 28- 
35 

carbon atoms, 33-35 
CO/C0 2 ratio, 32-33 
composition, 32-33 
during burning, 24 
quenching, 30-32 

reactions during cooling, 22, 28-30 
water gas reaction, 29, 32-33, 35 
Powder gas, erosive effects, 272-277, 
283-307 

see also Erosion products in gun bores 
apparatus for studying, 306-307 
carbon gases, 284-296 
cracking of bore surface, 274, 276- 
277 

gas leakage, 273-274, 562, 592-594 
gun liner joint, 274 
individual powder gas constituents, 
305-307 

mitigation, 300, 307, 325-327 
nitrogen, 300-304 
occlusion of gases, 274-275 
removal of material from bore, 272- 
273 

scouring action, 273 
sulfur gases, 271-272, 296-300 


CONFIDENTIAL 




INDEX 


669 


temperature and pressure, 274, 305- 
306 

Powder gas, pressure, 36-37, 41 
effect of pre-engraved projectiles, 
594-596 

effect on bore friction, 130 
erosive effects, 274, 305-306 
method of measuring, 80-81, 91-92 
peak pressure, 22 
quenching experiments, 31-32 
resistance gauge for measuring, 91-92 
starting time, 78 
Powder gas, radiation, 42-53 
absorption, 53 
characteristics, 45 
effect on burning rate, 27-28 
effect on gun temperature, 99 
peak radiation, 22 

photoelectric pyrometer, 46-47, 92, 
94 

spectral characteristics, 43-45 
starting time, 78 

temperature determinations from 
radiation measurements, 42-43 
Powder gas, temperature, 48-53 

composition as a function of tempera¬ 
ture, 37-38 

determination from radiation meas¬ 
urements, 42-43 

erosive effects, 251, 274, 305-306 
firings in a jet propulsion motor, 
51-52 

firings in closed vessels, 48 
firings in 3-in. gun, 48-51 
measurement, 46-47, 80, 82, 92, 94 
temperature gradient, 53 
Powder gas, thermodynamics, 35-42 
adiabatic flame temperature, 35, 
38-39 

change of internal energy with den¬ 
sity, 37 

energy released, 35-36 
enthalpy and entropy, 40 
equation of state, 36-38 
heat of explosion, 36, 41-42 
pressure, 36-37, 41 
specific heats, 40 
Powder metallurgy 
chromium, 345 
elkonite, 346 

molybdenum compounds, 346, 370- 
371 

Powders 

see also Powder gas 
adiabatic flame temperature, 38-39, 
312, 324-325 

characteristics of powders tested, 312 
density of loading, 63, 69, 71 
distribution of energy, 148-151 
equation of burning, 58 
explosives, 21 


FNH-M1; 122, 250, 270 
FNH-M2; 250, 270, 289 
granulation, 73-74 
impetus, 57, 73 
parameters, 61 
pyro powder, 323 
RDX, 289, 322, 327 
research recommendations, 327, 629- 
630 

Powders, burning process 
see Burning of powder 
Powders, composition 

changes in nitroglycerin content, 327 
changes to mitigate erosion, 283, 300, 
307, 325-327 
effect on burning rate, 27 
effect on gun erosion, 297, 299-300, 
307, 326-327 
ferrosilicons, 327 

Powders, double-base, erosiveness 
comparison with single-base, 308, 
326 

electron diffraction examinations of 
eroded specimens, 249 
erosion products, 248-251, 267 
FNH-M2 powder, 250, 270, 289 
liquefaction of bore surface, 266 
thermodynamic considerations, 250- 
251, 326 

X-ray examination of eroded speci¬ 
mens, 248-251, 310 

Powders, erosiveness, 288-289, 308-327 
black powder, 297, 299 
caliber .50 tests, 309-310, 313-322 
dependence on gas temperature, 251 
dimensional erosion rate vs flame 
temperature and heat input, 
324-325 

effect of composition, 251 
erosion rate, 324 

erosion vent plug tests, 310-311, 322- 
323 

heat input versus flame temperature, 

325 

mitigation, 283, 300, 307, 325-327 
pre-engraved projectiles, 600 
RDX powders, 322, 327 
thermal transformation vs. flame 
temperature, 325 
types of powder tested, 311-313 
Pow T ders, single-base, erosiveness 
carbon penetration into gun steel, 
288 

comparison with double-base, 308, 

326 

electron diffraction examination of 
eroded specimens, 249 
erosion products, 247-251, 267 
FNH-M1 powder, 250, 270 
liquefaction of bore surface, 266 
outer white layer, 247 


thermodynamic considerations, 250- 
251, 326 

X-ray examination of eroded speci¬ 
mens, 248-251, 310 

Pre-engraved projectiles, 523-533, 591- 
608 

accuracy, 597 
band material, 598-600 
bore friction, 145-147 
cadmium-plated, 599 
design, 14, 523-525 
double engraving, 595 
effect on powder gas pressure, 594- 
596 

erosion, 313-317, 595, 600-603, 606 
exterior ballistics, 528-532 
interior ballistics, 525-527 
manufacturing techniques, 529, 533 
parco-lubrized, 275, 599 
powders for, 600 
research recommendations, 630 
yaw, 526-527 

Pre-engraved projectiles, guns for fir¬ 
ing, 591-608 

.50 caliber, 592, 598-599, 601 
37-mm gun, T 47; 603-606 
90-mm gun, 606-608 
chromium-plated bores, 600-603 
Pre-engraved projectiles, loading and 
indexing mechanism, 551-556 
Browning caliber .50 machine gun, 
551-552 

Johnson 20-mm automatic aircraft 
guns, 552-553 

modification of guns, 551-552 
modified Bofors 40-mm mechanism, 
553, 555-556 

Pre-engraved projectiles, tests, 591-600 
effect of band material, 598-600 
effect of gas leakage on erosion of 
gun steel, 592-594 
effect of type of powder, 600 
effect on accuracy, 597 
erosion studies, 313-317 
methods of testing and measure¬ 
ment, 591-592 

pressure and velocity, 594-596 
T47 gun, 603-606 
Pressure in guns 

effect of gas leakage, 202 
effect of temperature, 122-123 
effect on burning rate of powder, 35 
gauges for measurement, 91-92 
in recoil cylinders, 82 
in water penetrated by a projectile, 
184-187 

powder gas, 31-32, 80-81, 91-92, 
305-306 

pre-engraved projectiles, 594-596 
pressure-time curves, 81, 136-137, 
145 


CONFIDENTIAL 



670 


INDEX 


pressure-travel curves, 63-65, 68, 
138-139, 607 

rotating gun-bands, 129-130, 152- 
162, 521-523 

Probert system of gun rifling and pro¬ 
jectile banding, 517 
Progressive stress-damage, 260, 279 
Projectile bands, rotating 
37-mm, 534-535 
75-mm, 535-536 
caliber .50 projectiles, 533-534 
design, 533-536 
effect of cadmium plating, 535 
engraving, 158-159 
erosion tests, 534 
low-stress types, 533-536 
melting, 100, 130, 135, 484 
pre-engraving; see Pre-engraved pro¬ 
jectiles 
shearing, 202 
Projectile engraving 
see also Pre-engraved projectiles 
bore friction, 129-130 
effect of Fisa protector, 611 
force required, 129, 533-534 
heat, 99-100, 318-319 
indication of stability, 465 
process, 158-159 
relation to erosion, 275 
relation to powder pressure, 594 
stresses, 275 
Projectiles 

see also Hypervelocity projectiles 
abrasive action, 275, 591 
acceleration, 79, 133-134 
center of pressure, 173 
characteristics, 176-179 
comparison of standard and hyperve¬ 
locity, 163 
deceleration, 79 
design, 518-536 

energy; see Energy of projectiles 
erosive effects, 275-276 
finned, 169 
head shapes, 176 
jump ratio, 526 
normal force, 174 
parco-lubrized, 275, 599 
research recommendations, 630 
rotating bands; see Projectile bands, 
rotating 

shatter, 189-191 

stability, 169-170, 176-177, 515, 
559-560 

trajectory, 63, 78-79, 83, 130, 133, 
177-179 

velocity, 91, 515, 594-596, 619- 
624 

yaw; see Yaw of projectiles 
Projectiles, stresses, 518-523 
axial stress, 523 


band pressure, 521-522 
British method of investigation, 522, 
536 

criterion of failure of a shell, 522- 
523 

detection of defects, 536-537 
external forces, 519 
formulas for shell design, 521-522 
high-explosive, 519 
hoop stress, 152,523 
limitations of theory, 520-521 
longitudinal stress, 153, 524-525 
method of calculating, 523 
Mises-Hencky theory, 522 
shear, 522 

skirted projectiles, 573-575 
stresses in base and walls, 520 
tangential stress, 515, 523 
types, 519-520 
yield strength, 523 
Projectiles, types 

armor-piercing, 9-10, 189-192, 557- 
558, 585-590, 626-628 
artillery-type for erosion-testing gun, 
313, 315-316, 595 
cadmium-plated, 354, 535, 599 
conical-headed, 175-177 
deformable, 169, 191 
nondeformable, 191-192 
pre-engraved, 313-317, 523-533, 
551-556, 591-608 

sabot, 169-170, 557-568, 585-587 
skirted, 569-590, 620 
spin-stabilized, 510, 515 
subcaliber, 165-170, 618-628 
tungsten carbide-cored; see Tungsten 
carbide-cored projectiles 
Propellants 
see Powders 

Protractor for differentiating curves, 
94-95 

Pulsers, electronic, 86 
Pump plating of gun barrels, 414 
Pyro powder, erosiveness of, 323 
Pyrolitic plating, 418-439 
see also Electroplated coatings 
carbonyl chemistry, 419-420 
chromium, 434-439 
molybdenum, 420-433 
refractory metals, 418 
tungsten, 433-434 
Pyrometer, photoelectric 
apparatus, 46-47 
design principles, 46 
radiation pyrometer, 92,94 
Pyrrohotite in eroded guns, 254 


Radial stress from projectile bands, 
152 

Radial-deflection scanner, 156-157 


Radiation 

effect on burning rate of powders, 27- 
28 

effect on cooling of guns, 126 
effect on heating of guns, 99 
from powder gas; see Powder gas, ra¬ 
diation 

Radiation pyrometer, 46-47, 92, 94 
Radioactive carbon, 33-35, 285 
Radioactive sulfur, experiments on ero¬ 
sion in guns, 296 

Radioactive tracers, gun erosion stud¬ 
ies, 240-243 

demonstration of equilibrium,240-241 
Geiger-Mueller counters, 242 
sampling, 241-242 
temperature of specimen, 242-243 
Range of hypervelocity projectiles, 9, 
164-165, 202 

R.D. system of gun rifling and projec¬ 
tile banding, 517 
RDX powder 
erosiveness, 322, 327 
reaction products, 289 
Recoil of gun 

see Gun recoil measurements 
Recoilmeters, 80, 86-87 
Recommendations for future research 
see Research recommendations 
Recording apparatus for ballistic meas¬ 
urements, 83-86 
oscillograph and recorder, 83-85 
oscillograph calibrators, 85-86 
pulsers, 86 

timing and developing of film, 86 
Refractaloy No. 70 
gun liners, 351-352 
thermochemical erosion resistance, 
407 

Research recommendations 
bore friction, 629 
drag coefficient, 176-177 
electroplated coatings, 417-41$ 
erosion in guns, 629 
gun liners, 368-369, 484, 629 
gun propellants, 629-630 
gun steel, 418, 482, 630 
hypervelocity guns, 628-630 
interior ballistics of guns, 628-629 
Johnson 20-mm gun, 548-549 
molybdenum carbonyl, 424 
pilot plants for chromium-plating of 
gun barrels, 499-500 
pre-engraved projectiles, 630 
sabot-projectiles, 630 
stability of projectiles, 176-177 
tapered-bore gun and skirted projec¬ 
tiles, 590 

vapor-phase plating, 432-433, 439 
water penetration by hypervelocity 
projectiles, 187 


CONFIDENTIAL 




INDEX 


671 


yaw of hypervelocity projectiles, 177 
Resistance gauge for powder-pressure 
measurements, 91-92 
Reynold's number, pre-engraved pro¬ 
jectiles, 529 

Rhodium, thermochemical erosion re¬ 
sistance, 355 
Rifling of guns, 515-517 
draw rifling, 504 

modification for Fisa protector, 611 
pre-engraved projectiles, 551-556 
Probert system, 517 
R.D. system, 517 

requirements of hypervelocity guns, 
515-516 

strength factors, 516-517 
swaging, 216-217, 275-276, 458-459, 
504 

torque on rifling, 516 
Rocket nozzles, molybdenum-plated, 
431 

Rotating projectile band 

see Projectile bands, rotating 
Rupture disks in erosion vent plugs, 
225, 228, 310 

Sabot-projectiles, 557-568 
ballistic limits, 585-587 
characteristics, 567 
comparison with skirted projectiles, 
586, 620 

conditions for optimum armor pene¬ 
tration, 170 
description, 557 
evaluation, 557-559 
interior ballistics, 75 
plastic, 562 
requirements, 557 

research recommendations, 568, 630 
stability, 169-170, 559-560 
time of flight, 165 

Sabot-projectiles, design, 15, 559-562 
materials, 562 
obturation, 562 
requirements, 559 
stability, 559-560 

Sabot-projectiles, release methods, 560- 
562 

axial release, 561, 565-567 
centrifugal release, 560-561, 563-565 
combined axial and centrifugal re¬ 
lease, 561-562, 564, 568 
gas pressure, 561 
Sabot-projectiles, sizes, 562-567 
20-mm gun, 562-563 

75- mm gun and howitzer, 563-564 

76- mm gun, 564-565 
90-mm gun, 564-567 
105-mm howitzer, 566 

St. Venant maximum strain theory, 
510 




Savage Arms Corporation, insertion of 
gun liners, 504 

Scanner, radial-deflection, 156-157 
Scoring by powder gas, 274 

75- mm guns 

band pressure measurements, 156- 
159 

projectile design, 535-536 
sabot-projectiles, 563-564 

76- mm guns, 122, 214, 564-565 
Shear-energy theory, 510 
Sheffield Precisionaire gauge, 487 
Shell stresses 

see Projectiles, stresses 
Shock waves, 83 
conical, 171 

in water penetrated by a bullet, 181— 
183, 187-188 

pressure as function of bullet velocity, 
186 

suppression, 576-578 
Silichrome gun steel, 337, 352-353, 483 
Silicon steel for gun liners, 353 
Silver, thermochemical erosion resist¬ 
ance, 354 

Single-base powders 

see Powders, single-base, erosiveness 
Skirted projectiles, 569-590 
advantages over sabot-type, 586, 620 
armor penetration tests, 585-590 
deformation, 573 
design, 572-573 
research recommendations, 590 
stresses, 573-575 
tungsten carbide core, 572 
use in 57/40-mm gun, 570-572, 584- 
590 

velocity limits, 576 
Skirted projectiles, tests, 579-585 
accuracy tests, 584-585 
cores, 583 

flight characteristics, 581-583 
front skirts, 581 
material, 581 
rear skirts, 580 
skirt position, 581 
windshield, 581, 583 
Sleeves for gun barrels, reinforcing, 
476-477 

see also Fisa protector for gun barrels 
Solenoids for determining projectile 
velocity, 91 
Specifications 

chromium-plated machine gun bar¬ 
rels, 413, 460 

Johnson 20-mm gun, 538-539 
nitrided machine gun barrels, 459- 
460 

stellite-lined gun barrels, 413 
Spectral characteristics of powder gas, 
43-45 


Sphalerite in eroded guns, 253 
Spin-stabilized projectiles, 515 
SR-4 strain gauge, 90 
Stability of projectiles 

see also Subcaliber projectiles, stabil¬ 
ity 

deformable projectiles, 169 
effect of yaw, 168 
effect on muzzle erosion, 12 
recommendations for future research, 
176-177 

sabot-projectile, 169-170, 559-560 
spin-stabilized, 515 
Star gauges, 196-198 
curves, 196-197, 467 
erosion tests, 222 
two- and three-point gauges, 197 
Stave-type gun liners 

see Molybdenum gun liners, stave- 
type 

Steel, diffusion of elements through, 
267-272, 284-296 
carbon, 264, 267-271, 284-296 
hydrogen, 274-275, 296 
nitrogen, 264, 267-269, 272, 300-304 
oxygen, 269-271 
sulfur, 296-300 
Steel, effect of powder gas 
see Powder gas, erosive effects 
Steel for gun barrels, 418, 480-484 
Chro-mow, 483 
critical temperature, 262-263 
Exelloy, 484 
HI A grade 160; 483 
isothermal transformation diagram, 
275 

microstructural constituents, 265 
molybdenum steels, 353, 381-382, 
482 

nickel and cobalt steels, 352 
nitralloy, 476 

nitrided steel, 353, 401-402, 459-460 
Peerless A, 483 

physical properties, 275,363,373,418 
Potomac die-steel, 470, 483-484 
research recommendations, 482, 630 
Silichrome, 337, 352, 483 
silicon, 353 
stainless, 351 
Templex, 482 
thermal expansion, 513 
Timken 1722A, 483 
TK hot die steel, 484 
tungsten steels, 352-353 
XB valve steel, 483-484 
Steel for projectiles, 191, 557, 581 
Steel gun liners, replaceable, 506-509 
design, 14, 507 
lined with stellite, 615 
90-mm liner, 507-509 
purpose, 506-507 


CONFIDENTIAL 



672 


INDEX 



Stellite, metallurgy and properties, 
391-397, 401-404 
availability, 396-397 
carbon content, effect on structure, 
394 

chemical composition, 392-393, 402, 
407 

decarburization, 402 
ductility, 392 
hardness, 395, 401 

heat treatment, effect on structure, 
394, 402 

hot working, 402 
machineability, 396 
mechanical properties, 396 
melting ranges, 331, 348, 396 
metallographic examination, 393- 
394 

nitriding, 401 

specific gravity and density, 395 
strength at high temperatures, 395 
thermal coefficients, 395-396 
thermochemical resistance, 391 
volume and dimensional stability, 
394-395 

work- and age-hardening, 394 
working properties for forging and 
rolling, 395 

X-ray examination, 394 
Stellite gun liners, 351, 397-406, 443- 
457 

coating methods, 397-398 
design, 398-399, 404-405, 443-446, 
503-509 

erosion, 124, 348-349 
for caliber .30 barrels, 454-456 
for caliber .50 barrels, 404, 446-454, 
473-484 

for caliber .60 barrels, 456-457 
for cannon, 405, 615-616 
infusion and incasting, 398, 404 
inmelting, 404 

insertion, 403, 443-446, 503-508 
integral chamber, 457 
length, 403, 473, 475 
liners in series, 404 
manufacturing methods, 451-454 
pilot plants, 451-453 
plating of barrels; see Stellite-lined 
and chromium-plated gun bar¬ 
rels 

salvage, 453 

sprayed coatings, 399-401 
steel barrels, special, 482-484 
stresses in, 155, 513 
thickness, 155, 403 
torch deposition, 399 
velocity-life, 473 

Stellite-lined and chromium-plated gun 
barrels, 473-484 

accuracy- and velocity-life, 473, 478 



barrel steel, 480-484 
caliber .30; 471-472 
CGL firing schedule, 474-475 
design, 477, 479, 484 
effects of barrel weight, 477-480 
external reinforcing sleeve, 476-477 
performance, 403-404, 465-469, 473- 
478 

plating procedure, 410-411, 414, 
486-491, 495-499 
specifications, 413, 459, 491 
tests, 473, 475-476 
Stereoscopic microscope, 236 
Strain gauges for measuring projectile 
displacement, 90-91 
Baldwin-Southwark gauge, 90 
SR-4; 90 

types of measurements, 82 
Strain measurements 

circumferential strain in gun barrel, 
155-156 

determination of bore friction, 132 
firing tests, 82, 106 
Strain-recording equipment, 161 
Stress, erosive effects, 230, 277, 279, 
510 

Stress in gun barrels, 510-515 

see also Band pressure of gun tubes 
caliber .50; 162 
circumferential, 155-156 
compressive, 511, 513 
design procedure for cannon tubes, 
511-512 
equations, 511 
equivalent stress, 511 
failure, 510-511 
integrated stresses, 153 
Lame formula, 152, 511 
land erosion, 510 
liner stresses, 512-515 
longitudinal, 153, 524-525 
Mises-Hencky theory, 510 
progressive stress-damage, 279 
radial stress, 152 
St. Venant theory, 510 
shrunk-in liners, 512 
stave-type liner, 515 
stress distribution in a thick-wall 
tube, 511 

tangential, 152, 511 
tensile, 511 

thermal stress, 154-156, 277, 279, 
513-514 

Stress in projectiles 
see Projectiles, stresses 
Stress measurements of projectile bands 
see Band pressure of gun tubes 
Stress raisers in guns, 279 
•Stressed erosion-vent apparatus, 230, 
272-275 

Subcaliber projectiles, 618-628 

UNCLAS 



see also Sabot-projectiles; Skirted pro¬ 
jectiles 

armor penetration, 192, 626-628 
ballistic coefficient, 165 
effectiveness of fire, 624-626 
flight characteristics, 165 
interior ballistics, 75 
military economics, 618-619 
pros and cons of subcaliber projectiles, 
619-620 

tungsten carbide cores, 165, 572, 583, 
585-587, 620 
velocity gains 619-624 
Subcaliber projectiles, stability, 165- 
170 

deformable projectiles, 169 
effect of yaw, 168 
formula, 168 

methods of obtaining stability, 168- 
169 

moment coefficient, 167 
over long arcs of the trajectory, 167 
projectile fired from a plane, 168 
sabot-projectiles, 169-170, 559-560 
theory, 165-166 

Subcaliber projectiles, velocity gains, 
619-624 

effect of gun caliber, 622 
effect of range, 622-624 
limiting effect of powder gas kinetic 
energy, 620-622 
Sulfides 

copper and zinc, 252-253 
hydrogen, 271-272, 292-294, 296, 
427 

identification in eroded guns, 240, 
252-253, 298-299 

Sulfur 

compounds, 252-253, 299 
effect on erosion, 240-241, 252, 296, 
299 

penetration into steel, 299 
radioactive, 296 

reaction with carbon monoxide, 291- 
293 

reaction with gun steel, 296-300 
Sulfur prints, 298-299 
Supersonic cone, motion of, 170-179 
Swaging of rifling, 216-217, 275-276, 
458-459, 504 

Tangential resistance in gun tubes, 511 
Tangential stress 
projectiles, 523 
stave-type liners, 515 
Tantalum 

electrodeposition from fused salt 
baths, 409 
gun liners, 347 
Tapered-bore guns, 569-590 
adaptors, 571, 575-576, 579-580 






design, 15, 575-576 
57/40-mm, 570-572, 580, 584-590 
interior ballistics, 75 
suggestions for future development, 
590 

tests, 578-580 
venting, 576-578 
wear of gun, 579-580 
Target patterns, moving, 448-449 
Temperature, effect on gun performance 
see Gun firing, thermal effects 
Temperature-indicating paints, 106- 
107 

Temperature-time curves for machine 
gun barrels, 116-117, 461 
Templex gun steel, advantages, 482 
Terminal ballistics of hypervelocity 
projectiles, 180-192 
armor perforation, 189-192 
disruption of a liquid by projectile, 
•180-189 
TEW steel alloy 
gun liners, 352 

resistance to surface cracking, 335 
thermochemical erosion resistance, 
407 

Thermal alteration of bore surface 
see Bore surface, thermal transforma¬ 
tion 

Thermal analyzer, 107 
Thermal effects of gun firing 
see Gun firing, thermal effects 
Thermal shock, factor in gun erosion, 
261, 277 

Thermal softening of gun bores, 261- 
263 

development of martensite, 263 
heating process, 261-262 
mushrooming, 263 
single-shot guns, 261-262 
Thermal stresses in gun barrels, 154- 
156 

circumferential strain at outer sur¬ 
face, 155-156 

cracking of bore surface, 277, 279 
liner insertion, 513-514 
stresses at a point, 154-155 
Thermindex paint E6; 107 
Thermochemical erosion resistance of 
metals and alloys, 343-355 
aluminum, 354 
beryllium, 354 

chromium-base alloys, 406-407 
cobalt, 347-348 
columbium, 354 
gold-palladium, 354-355 
hastelloys, 407 
iridium, 355 

iron-base alloys, 406-407 
nickel-base alloys, 406-407 
Refractaloy No. 70; 407 


INDEX 


MlLAMlNtl 


673 


rhodium, 355 
silver, 354 

stellite No. 6; 391-392 
stellite No. 21; 391 
TEW steel alloy, 407 
titanium, 354 
zinc, 354 

Thermocouples for measuring gun tem¬ 
perature, 105-106, 113 
Thermodynamics of powder gas 
see Powder gas, thermodynamics 
37-mm gun, T47 
ballistic firings, 78 
bore friction measurements, 144-148 
burning-rate constants for powders, 
27 

design, 603-605 

energy distribution of gun powder, 
151 

erosion test with pre-engraved pro¬ 
jectiles, 606 

heat input to bore surface, 101 
pressure-time curves, 145 
37-mm guns 

design of projectiles, 534-535 
tests with Fisa protector, 614 
37-mm guns, band pressure measure¬ 
ments, 156-159, 161-162 
apparatus and method, 156-157 
band pressure per inch circumference, 
157 

effective area of interference (EAI), 
157 

friction coefficient, 157 
prediction of band pressures, 157 
pressure-time curves, 145 
process of engraving, 158-159 
radial deformation, 156 
Thorium oxide refractories for melting 
molybdenum, 387 

Thread-collocating gauge for machining 
gun tubes, 578 
3-in. gun 

antiaircraft, 124 

bore friction measurements, 135-144 
energy balance during firing, 149- 
151 

liners, 249, 616-617 
Mark 3; 77-78, 101 
powder gas temperature measure¬ 
ments, 48-51 

pressure and friction curves, 131 
pressure-time curves, 81, 136-137 
pressure-travel curves, 68, 138-139 
70 caliber, 119-120 

3-in. gun, band pressure measurements, 
159-161 

axial strain, 154 

graphical determination, 160-161 
tangential strain, 153 
3.7-in. gun, 119, 121, 249 


Ti-Kovar (alloy) for gun tube sleeves, 
514 

Timing-pulse generator, 86 
Timken 1722A gun steel, performance, 
483 

Titanium, thermochemical erosion re¬ 
sistance, 354 
TK hot die steel, 484 
Tracer isotopes, gun erosion studies, 
240-243 

demonstration of equilibrium, 240- 
241 

Geiger-Mueller counters, 242 
sampling, 241-242 
temperature of specimen, 242-243 
Tracer photography, 83, 177-179 
Trajectory of projectiles, determina¬ 
tion from photography, 83, 177- 
179 

Transformation of surface layers 

see Bore surface, thermal transfor¬ 
mation 

Trivalent chromium, 412 

Troostite band in gun bore surfaces, 247 

Tube stresses 

see Stress in gun barrels 
Tungsten 

alloys, 376, 417, 434 
carbonyl, 433 
chloride, 433 

gun liners, 346-347, 352-353 
vapor-phase plating, 433-434 
Tungsten carbide-cored projectiles 
armor perforation, 190, 192, 628 
skirted projectiles, 572 
strength, 583 

subcaliber projectiles, 165, 620, 628 
Tungsten Electrodeposit Corporation, 
tungsten alloy deposition, 417 
Turbulent flow, 53, 266, 273, 613 
20-mm guns 

Hispano-Suiza, AN-M2; 118, 121, 
123, 133 

Johnson, 538-549 
projectiles, 171, 562-563 

University of California, thermal ana¬ 
lyzer, 107 

Vapor-phase plating, 418-439 
see also Electroplated coatings 
carbonyl chemistry, 419-420 
chromium, 434-439 
molybdenum, 420-433 
refractory metals, 418 
research recommendations, 432-433, 
439 

tungsten, 433-434 
Velocimeters, 86-90 

differentiators for, 89-90 
linear, 89 







measurement of gun velocity in recoil, 
80, 88-89 
rotary, 88-89 
Velocity, muzzle 
see Muzzle velocity 

Velocity measurements on gun barrels 
see also Velocity-life of caliber .50 
gun barrels 

effect of chromium-plated bores, 458, 
463, 469, 602-603 
gun in recoil, 80 

velocity-drop, 202, 508, 595-596 
Velocity of projectiles 
effect on armor perforation, 192 
means of increasing, 66, 549-550, 
619-620 

pre-engraved projectiles, 594-596 
solenoids for determining, 91 
spin-stabilized, 515 
subcaliber, 619-624 
Velocity-life of caliber .50 gun bar¬ 
rels 

chromium-plated, 458, 466-468, 473, 
478 

stellite-lined, 450-451, 473, 478 
Vent plugs, erosion 
see Erosion vent-plug tests 
Vinylite coatings for studying ero,ded 
gun bore surfaces, 236 


Water ballistic pendulum, 184-185 
Water gas reaction, 29, 32-33, 35, 269- 
271 

Water-cooled guns 
cook-off tests, 121 
disadvantages, 127 
water jackets, 127 
Weapons, gas-operated 
see Johnson 20-mm gun 
White layer on eroded gun bores 
austenite, 247, 264 
carbon, 269 

removal technique, 256 
single-base powders, 247 
Windshields for projectiles, 583 
Wire-in-bore technique of interior bal¬ 
listic measurements, 95-96 
Wiistite in eroded gun, 249, 253 

XB valve steel for gun barrels, 483- 
484 

X-ray diffraction in gun erosion stud¬ 
ies, 237-239, 248-251 
bore surfaces, 248-249 
carbon monoxide reactions, 291 
comparison with electron diffraction, 
237 

erosion-testing gun barrels, 251, 310, 
321 


fired metal particles, 249-250 
identification of reaction products, 
247-249, 291, 301 
Stellite No. 21; 394 
test blocks exposed to powder gases, 
249 

unidentified lines, 239, 254 
use of different wavelengths, 238 
vent plugs, 228, 230 


Yaw of projectiles 
effect of muzzle erosion, 211 
effect on drag coeciffient, 175-176 
effect on hypervelocity projectiles, 
170-175 

effect on stability, 168, 463 
pre-engraved, 526-527 
research recommendations, 177 
within gun bore, 526-527 


Zinc, thermochemical erosion resist¬ 
ance, 354 

Zinc sulfides (gun erosion product), 
252 

Zirconium, deoxidizer for chromium, 
359-360 

Zirconium-nickel gun liners, 350 



























































































































































































































































































































































































































































































































































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